cone penetration test design guide for state geotechnical
TRANSCRIPT
Cone Penetration TestDesign Guide for
State Geotechnical Engineers
Author David Saftner Report Number 2018-32
Date Published November 2018
Minnesota Department of Transportation Research Services amp Library
395 John Ireland Boulevard MS 330 St Paul Minnesota 55155-1899
mndotgovresearch
To request this document in an alternative format such as braille or large print call 651-366-4718 or 1-800-657-3774 (Greater Minnesota) or email your request to ADArequestdotstatemnus Please request at least one week in advance
Technical Report Documentation Page
1 Report No
MNRC 2018-32
2 3 Recipients Accession No
4 Title and Subtitle
Cone Penetration Test Design Guide for State Geotechnical
Engineers
5 Report Date
November 2018 6
7 Author(s)
Ryan Dagger David Saftner and Paul Mayne
8 Performing Organization Report No
9 Performing Organization Name and Address
Department of Civil Engineering University of Minnesota Duluth 1405 University Drive Duluth MN 55812
10 ProjectTaskWork Unit No
CTS2017022 11 Contract (C) or Grant (G) No
(C) 99008 (WO) 249
12 Sponsoring Organization Name and Address
Minnesota Department of Transportation Research Services amp Library 395 John Ireland Boulevard MS 330 St Paul Minnesota 55155-1899
13 Type of Report and Period Covered
Final Report 14 Sponsoring Agency Code
15 Supplementary Notes
httpmndotgovresearchreports2018201832pdf 16 Abstract (Limit 250 words)
The objectives of this project are focused on a new cone penetration testing (CPT) geotechnical design manual for highway and transportation applications based on recent research and innovation covering the period from 2000 to 2018 A step-by-step procedure is outlined on how to use CPT data in the analysis and design of common geotechnical tasks Previous manuals are either very outdated with information from 1970-1996 or not appropriately targeted to transportation works This design document introduces modern and recent advancements in CPT research not otherwise captured in legacy manuals from the 1990rsquos and earlier Examples and case studies are provided for each topic interpreted using CPT measures In the manual a step-by-step procedure is outlined on how to use CPT data in analysis and design for typical geotechnical practices These topics which are applicable both to state highways and local roads include bridge foundations (including shallow footings and deep foundations) and soil characterization (including determination of standard soil engineering properties)
17 Document AnalysisDescriptors
Cone penetrometers Geographic information systems Soils by
properties Bridge foundations Soil mechanics
18 Availability Statement
No restrictions Document available from
National Technical Information Services
Alexandria Virginia 22312
19 Security Class (this report)
Unclassified
20 Security Class (this page)
Unclassified
21 No of Pages
225
22 Price
CONE PENETRATION TEST DESIGN GUIDE FOR STATE
GEOTECHNICAL ENGINEERS
Prepared by
Ryan Dagger
David Saftner
Department of Civil Engineering
University of Minnesota Duluth
Paul W Mayne
School of Civil amp Environmental Engineering
Georgia Institute of Technology Atlanta
November 2018
Published by
Minnesota Department of Transportation
Research Services amp Library
395 John Ireland Boulevard MS 330
St Paul Minnesota 55155-1899
This report represents the results of research conducted by the authors and does not necessarily represent the views or policies
of the Minnesota Department of Transportation the University of Minnesota Duluth or the Georgia Institute of Technology
This report does not contain a standard or specified technique
The authors the Minnesota Department of Transportation the University of Minnesota Duluth and the Georgia Institute of
Technology do not endorse products or manufacturers Trade or manufacturersrsquo names appear herein solely because they are
considered essential to this report because they are considered essential to this report
TABLE OF CONTENTS
CHAPTER 1 INTRODUCTION1
CHAPTER 2 DIRECT CPT METHOD FOR SOIL CHARACTERIZATION 4
21 Introduction 4
22 Soil Unit Weight 6
23 CPT Material Index 7
231 Step 1 Normalized Sleeve Friction 7
232 Step 2 Iteration 8
24 Soil Behavior Type (SBT) 8
241 Step 1 8
242 Step 2 9
25 Effective Stress Friction Angle 9
26 Stress History 10
27 Lateral Stress Coefficient 11
28 Undrained Shear Strength 12
29 Ground Stiffness and Soil Moduli 12
210 Coefficient of Consolidation 13
211 Hydraulic Conductivity14
212 Example Problems 15
2121 Example 1 Direct CPT Methods for Geoparameters on Sands 15
2122 Example 2 Direct CPT Methods for Geoparameters on Clay 33
CHAPTER 3 DIRECT CPT METHOD FOR SHALLOW FOUNDATIONS 56
31 Procedure 56
311 Step 1 Estimating Footing Dimensions57
312 Step 2 Soil Characterization 58
313 Step 2a Foundation soil formation parameter59
314 Step 3 Soil elastic modulus and Poissonrsquos ratio 60
315 Step 4 Net cone tip resistance 60
316 Step 5 Bearing capacity of the soil 61
317 Step 6 Settlement61
318 Step 7 Final Check 61
32 Example Problems 62
321 Example 3 Direct CPT Method on Sands 62
CHAPTER 4 DIRECT CPT METHOD FOR DEEP FOUNDATIONS 70
41 Introduction70
42 Modified UniCone Method72
421 Step 172
422 Step 272
423 Step 373
424 Step 473
43 Axial Pile Displacements 73
44 Example Problems 74
441 Example 5 Direct CPT Methods Axial Pile Capacity74
REFERENCES 87
APPENDIX A
APPENDIX B
APPENDIX C
1
2
3
4
5
6
7
8
9
10
11
12
13
14
15
16
17
18
19
20
21
22
23
24
LIST OF FIGURES
Figure Geoparameters determined from CPT 1
Figure Conventional method for shallow foundation design compared to direct CPT method 2
Figure Direct CPT evaluation of axial pile capacity 3
Figure Soil unit weight from CPT sleeve friction 6
Figure Approximate ldquorules of thumbrdquo method using CPT sounding from Wakota Bridge MN 7
Figure SBT zones using CPT Ic 9
Figure Yield stress exponent compared to CPT material index 11
Figure k vs dissipation time for 50 consolidation (Mayne 2017) 15
Figure CPT data from Benton County Minnesota for example problem 1 16
Figure Soil layers using ldquorules of thumbrdquo17
Figure Soil layer 1 using SBT method 23
Figure Soil layer 2 using SBT method 24
Figure Soil layer 3 using SBT method 25
Figure Soil layer 4 using SBT method 26
Figure CPT data from Minnesota for example problem 2 34
Figure Soil layers using ldquorules of thumbrdquo36
Figure Soil layer 1 using SBT method 42
Figure Soil layer 2 using SBT method 43
Figure Soil layer 3 using SBT method 44
Figure Soil layer 4 using SBT method 45
Figure Dissipation t50 data 53
Figure Direct CPT method for shallow foundations56
Figure Direct CPT method introduction 58
Figure Differentiation of porewater pressure measurement locations (Lunne et al 1997) 58
Figure 25 Foundation soil formation parameter hs versus CPT material index Ic (Mayne 2017) 60
Figure 26 Diagram of footing profiles for Example 362
Figure 27 CPT data from Northern Minnesota 63
Figure 28 Schematic of foundation design associated with CPT data 64
Figure 29 Soil type for example problem 367
Figure 30 Direct CPT evaluation of axial pile capacity 70
Figure 31 UniCone Method soil behavior type using CPT (Mayne 2017) 71
Figure 32 Modified UniCone Method soil behavior type using CPT (Mayne 2017) 71
Figure 33 Deep foundation end bearing pile diagram74
Figure 34 CPT data from Minnesota for example problem 5 75
Figure 35 Soil layers using ldquorules of thumbrdquo for pile capacity example using direct CPT method 76
Figure 36 Soil layer 1 using SBT method 81
Figure 37 Soil layer 2 using SBT method 82
Figure 38 Soil layer 3 using SBT method 83
Figure 39 Soil layering compared to pilings 85
LIST OF TABLES
Table 1 Geoparameters calculated directly from CPT 4
Table 2 Minor Geoparameters 5
Table 3 Soil type compared to exponent m 10
Table 4 Geoparameters evaluated for Case Example 117
Table 5 Geoparameters 35
Table 6 Bearing capacity defined by soil type61
Table 7 SCPT Results 62
CHAPTER 1 INTRODUCTION
The purpose of this manual is to provide an analysis of soil characterization shallow footings and deep
foundations using direct cone penetration testing (CPT) methods Geotechnical site characterization is
important for evaluating soil parameters that will be used in the analysis and design of foundations
retaining walls embankments and situations involving slope stability Common practice is to determine
these soil parameters through conventional lab and in-situ testing An alternative method uses CPT
readings of cone tip resistance (qt) sleeve friction ( fs) and pore pressure (u2) directly to determine these
parameters such as unit weight effective friction angle undrained shear strength and many others
(Figure 1) Direct CPT methods are provided for these parameters which will be used in the designs for
shallow and deep foundations
Figure 1 Geoparameters determined from CPT
Designing shallow foundations is typically done in a two-part process determining the bearing capacity
and expected settlement (commonly referred to as displacement) of the soil to approximate the
required size and shape of a foundation The older traditional methods are no longer required with
many approaches existing for using CPT directly in the design of shallow foundations Results from these
methods can provide a direct assessment of bearing capacity andor settlement A specific approach to
the direct method has been recommended and tested using a database of 166 full-scale field load tests
(Figure 2) A one-part process is used to scale the measured cone penetrometer readings (ie
measured cone tip resistance sleeve friction and pore water pressure) to obtain the bearing capacity
and settlement of the soil A step-by-step procedure has been created to transition from the CPT data to
bearing capacity with settlement accounted for The steps consist of estimating a design footing width
1
and length while using the process in the Soil Characterization section to determine soil parameters
directly from the CPT data
Figure 2 Conventional method for shallow foundation design compared to direct CPT method
There are upwards of 40 different direct CPT methods that have been developed over the past five
decades to determine the axial compression capacity of a piling foundation Earlier direct CPT methods
relied on hand-recorded information where mechanical-type CPT cone tip resistance data would be
collected at 20 cm intervals
Herein the method recommended for deep foundation design is the Modified UniCone method which
uses all three readings of the modern electronic piezocone penetrometer (CPTu) while addressing a
variety of pile foundation types (Figure 3) The modified UniCone Method is based on a total of 330 pile
load tests (three times the original UniCone database) that were associated with SCPTu data Two
computer software programs have been recommended for analyzing pile movements andor
settlements
2
Figure 3 Direct CPT evaluation of axial pile capacity
3
Symbol Parameter
γt Soil total unit weight
Ic CPT material index
SBT Soil behavior type (SBT)
ʹ σp Preconsolidation stress
YSR Yield stress ratio
ϕʹ Effective friction angle
Ko Lateral stress coefficient
su Undrained shear strength
Dʹ Constrained modulus
Eʹ Drained Youngrsquos modulus
CHAPTER 2 DIRECT CPT METHOD FOR SOIL
CHARACTERIZATION
21 INTRODUCTION
A direct CPT method for determining the value of each of various geoparameters shown in Table 1 is
provided The parameter Ic is used in the derivation of several sequential parameters This section of
the guide may be referred to as these parameters are used in calculations throughout the shallow
foundations and deep foundations sections
Table 1 Geoparameters calculated directly from CPT
4
Kʹ Bulk modulus
ks Subgrade reaction modulus
MR Resilient modulus
Gmax Small-strain shear modulus
k Coefficient of permeability
cv Coefficient of consolidation
Symbol Parameter Equation
ρt Mass Density ρt = γtga
where ga = 98 ms2
σvo Total Stress σvo asymp Σ (γti middot Δzi)
c Effective cohesion ʹ Empirical c asymp 003σp
In clays c asymp 01cu
Other minor parameters can be used in calculations of the geoparameters in Table 1 These minor parameters are provided in Table 2
Table 2 Minor Geoparameters
5
22 SOIL UNIT WEIGHT
The total soil unit weight can be estimated from CPT sleeve friction resistance as shown in Figure 4
(Mayne 2014) This method is not applicable to organic clays diatomaceous soils peats or sensitive
soils
Figure 4 Soil unit weight from CPT sleeve friction
Use Equation 1 to calculate the soils total unit weight
1 120574119905 = 120574119908 ∙ [122 + 015 ∙ ln (100 ∙ 119891119904 + 001)]
120590119886119905119898
Since soil unit weight is required for determining most geoparameters (including Soil Behavior Type)
estimating the soil type of the layers using the ldquorules of thumbrdquo method is a good first step before
determining more precise layering (Figure 5) Once a unit weight is determined for each noticeable layer
change these results can be used in later calculations such as ldquoCPT Material Indexrdquo To use the ldquorules
of thumbrdquo method some helpful guidelines are to assume sands are identified when qt gt 725 psi and u2
asymp uo while the presence of intact clays are prevalent when qt lt 725 psi and u2 gt uo The magnitude of
porewater pressures help to indicate intact clays such as soft (u2 asymp 2middotuo) firm (u2 asymp 4middotuo) stiff (u2 asymp 8middotuo)
and hard (u2 asymp 20middotuo) Fissured overconsolidated clays tend to have negative u2 values such that u2 lt 0
6
Figure 5 Approximate ldquorules of thumbrdquo method using CPT sounding from Wakota Bridge MN
23 CPT MATERIAL INDEX
The development of the CPT material index Ic has improved the initial classification of soil types and
calculation of soil parameters as shown in Table 1 To calculate Ic follow steps 1 and 2
231 Step 1 Normalized Sleeve Friction
Calculate Fr using Equation 2 if not provided with the CPT data gathered
2
7
119891119904 119865119903() = 100 ∙
(119902119905minus120590119907119900)
232 Step 2 Iteration
Iterate using Equations 3-5 to determine Ic by initially using n = 1 to calculate a starting value of Ic The
exponent n is soil-type dependent n = 1 (clays) n asymp 075 (silts) and n asymp 05 (sands) Iteration converges
quickly which is generally after the 3rd cycle
3 (119954119957minus120648119959119952)120648119938119957119950 = 119951 119928119957119951 prime (120648119959119952120648119938119957119950)
4
prime 120590119907119900 119899 = 0381 ∙ 119868119888 + 005 ( ) minus 015 120590119886119905119898
119899 le 10
5 119920119940 = radic[120785 120786120789 minus 119949119952119944119928119957119951]120784 + [120783 120784120784 + 119949119952119944119917119955]120784
24 SOIL BEHAVIOR TYPE (SBT)
Typically soil samples are not taken when using CPT The soil types are then inferred from the qt fs and
u2 readings To determine the types of soil from CPT data follow steps 1 and 2
241 Step 1
To determine the soil layers from CPT results calculate Ic by following the steps under section ldquoCPT
Material Indexrdquo After Ic has been determined through all specified depths use Figure 6 to classify the
type of soil by comparing each Ic value to normalized CPT readings (Fr and Qtn) from Equations 2 and 3
8
Figure 6 SBT zones using CPT Ic
242 Step 2
To determine if any of the soil layers contain ldquosensitive clays and siltsrdquo from zone 1 or ldquovery stiff
overconsolidated (OC) soilrdquo from zones 8 and 9 use Equations 6 and 7 If any soil layers are found within
zone 1 by Equation 6 then caution should be taken as these clays are prone to instability collapses and
difficulties in construction performance Very stiff OC sands to clayey sands of zone 8 (15 lt Fr lt 45)
and very stiff OC clays to silts of zone 9 (Fr gt 45) can be identified by Equation 7
6 Equation 6 errata This is an exponential expression (see Figure 5)
119876119905119899 lt 12119890119909119901(minus14 ∙ 119865119903)
7 119876119905119899 gt 0005(119865119903minus1)minus00003(119865119903minus1)2minus0002
1
Errata See terms and coefficients in Figure 5 above (numbers cannot be rounded off)
After each CPT reading has been assigned a zone from Figure 14 a visual representation can be made to
show the predominant layers by soil types (Figure A5)
25 EFFECTIVE STRESS FRICTION ANGLE
The effective friction angle (ϕ) is used to govern the strength for sands and clays where Equation 8 is
used for sands and Equation 9 is used for clays The value of ϕ for sands is derived from Ic so before ϕ
can be calculated refer to the iteration of Qtn under the ldquoCPT Material Indexrdquo section Once the values
9
Soil Type m
Fissured clays 11
Intact clays 10
Sensitive clays 09
Silt mixtures 085
Silty sands 080
Clean sands 072
Note m may be higher than 11 in fissured clays
of Ic and Qtn are calculated the type of soil can be determined from the section ldquoSoil Behavior Type
SBTrdquo The type of soil will dictate which equation to use for ϕ
Sands
8 120601prime(deg) = 176deg + 110deg log(119876119905119899)
Clays
9 119902119905minus120590119907119900 120601prime(deg) = 295deg ∙ 119861119902
0121 ∙ [0256 + 0336 ∙ 119861119902 + log ( )] prime 120590119907119900
Where = 119861119902 (1199062 minus 119906119900)(119902119905 minus 120590119907119900)
26 STRESS HISTORY
Determining the stress history can be characterized by an apparent yield stress ratio of the form
10 prime 120590119901
119884119878119877 = prime 120590119907119900
Where σp is defined as the preconsolidation stress or effective yield stress (Equation 10) The YSR is the
same equation as the more common overconsolidation ratio (OCR) but is now generalized to
accommodate mechanisms of preconsolidation such as ageing desiccation repeated cycles of wetting-
drying repeated freeze thaw cycles and other factors
11 prime prime 120590119901 = 120590119910 = 033(119902119905 minus 120590119907119900)119898prime
Where σp σy σvo and qt have units of kPa and the value of m depends on soil type with typical values shown in Table 3
Table 3 Soil type compared to exponent m
10
The value of m for non-fissured soils and inorganic clays and silts is derived from Ic (Figure 7) so before
m can be calculated (Equation 12) refer to the iteration of Ic under the ldquoCPT Material Indexrdquo section
Determine Ic for all soil layers before calculating m
12 028
119898prime = 1 minus 25 119868119888 1+( frasl ) 265
Once m is known for all soil layers the YSR can be determined
Figure 7 Yield stress exponent compared to CPT material index
A limiting value of YSR can be reached for clays and sands It can be calculated using Equation 13
13 (1 sin(emptyprime))
(1+sin(emptyprime)) 119884119878119877119897119894119898119894119905 = [ ]
(1minussin(emptyprime))2
27 LATERAL STRESS COEFFICIENT
The lateral stress coefficient Ko = σhoσvo commonly referred to as the at-rest condition is used to
represent the horizontal geostatic state of soil stress Ko can be calculated using Equation 14 but
11
14
119870119900 = (1 minus sin(emptyprime)) ∙ 119884119878119877sin( emptyprime)
sections such as ldquoStress Historyrdquo and ldquoEffective Stress Friction Anglerdquo will need to be referred to for
determining parameters ϕ and YSR
A maximum value for Ko can be determined by Equation 15
15 (1+sin(emptyprime))
119870119900119898119886119909 = = 1199051198861198992(45deg + emptyprime2) (1minussin(emptyprime))
28 UNDRAINED SHEAR STRENGTH
Loading on soils can result in fully drained partially drained or fully undrained conditions Sands
typically produce drained cases due to their high permeability but exceptions may occur in loose sands
during fast loading where the water does not have sufficient time to dissipate Clays exhibit low
permeability and thus often result in undrained loading cases when a load is applied quickly For soft-
firm clays the undrained shear strength (su) can be determined from CPT via Equation 16 where the
value of the bearing factor Nkt can be taken as 12
16 119902119905minus120590119907119900 119904119906 =
119873119896119905
In the case of remolded undrained shear strength from CPT su asymp fs
29 GROUND STIFFNESS AND SOIL MODULI
Determining the grounds stiffness can be measured from geoparameters such as the constrained
modulus (D) drained Youngrsquos modulus (E) bulk modulus (K) subgrade reaction modulus (ks) resilient
modulus (MR) and small-strain shear modulus (Gmax)
17 119863 prime asymp 5 ∙ (119902119905 minus 120590119907119900)
18
12
119864 prime = 119863 prime
11
19 119864prime
119870prime = [3∙(1minus2119907prime)]
The subgrade modulus (ks) is a combination of soil-structural properties which creates a parameter that
depends on the ground stiffness and the size of the loaded element
20 119864prime
119896119904 = [119889∙(1minus1199072)]
The resilient modulus MR applies to pavement analysis and design and can be calculated using Equation
21 where qt and fs are in MPa
21 053 119872119877 = (146119902119905 + 135511989111990414 + 236)244
The small strain shear modulus Gmax is a representation of the initial stiffness of all soils and rocks
Graphically it is the beginning portion of all stress-strain-strength curves for geomaterials Use Equation
22 to determine Gmax
22 119866119898119886119909 = 120588119905 ∙ 1198811199042
Where Vs = shear wave velocity (Equation 23) as measured by seismic cone penetration tests (SCPT) If
only cone penetration tests (CPT) or piezocone (CPTu) data are available the shear wave velocity may
be estimated from
23 03 119891119904 119881119904 (119898frasl119904) = [101 ∙ log(119902119905) minus 114]167 ∙ (100 ∙ )
119902119905
Where qt and fs have units of (kPa)
210 COEFFICIENT OF CONSOLIDATION
The coefficient of consolidation (cv) controls the rate that foundation and embankment settlements
occur By using results of CPT dissipation tests that measure the change in u2 readings over time the
value of cv can be determined Using Equation 24 cv can be determined based on piezocone dissipation
13
curves The equation below requires an estimate of the in-situ rigidity index (IR) of the soil If results of
SCPTU are available then IR may be determined from Gmax and qt per equation A38
24 0030∙(119886119888)2∙(119868119877)075
119888119907 = 11990550
Where ac = penetrometer radius (178 cm for 10-cm2 cone 220 cm for 15-cm2 cone) t50 = time to reach 50 dissipation IR = Gsu = undrained rigidity index G can be determined from the ldquoGround Stiffness and Soil Modulirdquo section
211 HYDRAULIC CONDUCTIVITY
The hydraulic conductivity also known as the coefficient of permeability (k) expresses the flow
characteristics of soils and has units of cms or feetday One method of calculation would be to use
Equation 25 where cv would need to be determined from ldquoCoefficient of Consolidationrdquo and D would
need to be determined from ldquoGround Stiffness and Soil Modulirdquo
25 119888119907∙120574119908 119896 =
119863prime
An alternative approach developed for soft normally-consolidated soils (Figure 8) is shown in Equation
26 where t50 (sec) values are used directly in assessing k in (cms)
26
14
125
119896 asymp ( 1
) 251∙11990550
Figure 8 k vs dissipation time for 50 consolidation (Mayne 2017)
212 EXAMPLE PROBLEMS
2121 Example 1 Direct CPT Methods for Geoparameters on Sands
Several geoparameters need to be determined based on the given CPT data collected for the South
Abutment of a bridge in Benton County MN (Figure 9) The groundwater table (GWT) was measured at
17 feet Determine all the geoparameters found in Table 4 at depths of 0 feet to 30 feet All sand layers
can be assumed ldquodrainedrdquo with ν = 02
15
Figure 9 CPT data from Benton County Minnesota for example problem 1
16
Table 4 Geoparameters evaluated for Case Example 1
Symbol Parameter Symbol Parameter
γt Soil total unit weight Ko Lateral stress coefficient
Ic CPT material index Dʹ Constrained modulus
SBT Soil behavior type (SBT) Eʹ Drained Youngrsquos modulus ʹ σp Preconsolidation stress Kʹ Bulk modulus
YSR Yield stress ratio MR Resilient modulus
ϕʹ Effective friction angle Gmax Small-strain shear modulus
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 10 Soil layers using ldquorules of thumbrdquo
γw = 6224 pcf
17
Layer 1
From Figure 10 fs = 17 psi taken as a representative value of the layer
17 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1204 pcf
145 psi
Layer 2
From Figure 10 fs = 17 psi taken as a representative value of the layer
17psi
γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1204 pcf 145 psi
Layer 3
From Figure 10 fs = 12 psi taken as a representative value of the layer
12 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 4
From Figure 10 fs = 7 psi taken as a representative value of the layer
CPT Material Index
7 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1121 pcf
145 psi
Layer 1
From Figure 10 fs = 17 psi and qc = 3500 psi
18
qt = qc + u2 ∙ (1 minus a) = 3500 psi + 3 psi ∙ (1 minus 08) = 3501 psi
= γt ∙ 6 feet = 1204 pcf ∙ 6 feet = 7225 psf = 50 psi σvo
fs 17 psi Fr() = 100 ∙ = 100 ∙ = 049
(qt minus σvo) (3501 psi minus 50 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 5 psi minus 0 psi = 50 psi
(qt minus σvo)σatm (3501 psi minus 50 psi )145 psi = = Qtn prime )n (σvoσatm (50 psi145 psi)n
prime σvo 50 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(049)]2
Qtn = 35329 n = 036 lt 10 Ic = 13
Layer 2
From Figure 10 fs = 17 psi and qc = 3500 psi
19
qt = qc + u2 ∙ (1 minus a) = 3500 psi + 0 psi ∙ (1 minus 08) = 3500 psi
= 7225 psf + γt ∙ 6 feet = 7225 psf + 1204 pcf ∙ 6 feet = 1445 psf = 100 psi σvo
fs 17 psi Fr() = 100 ∙ = 100 ∙ = 049
(qt minus σvo) (3500 psi minus 100 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 100 psi minus 0 psi = 100 psi
(qt minus σvo)σatm (3500 psi minus 100 psi )145 psi = = Qtn prime )n (σvoσatm (100 psi145 psi)n
prime σvo 100 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
47 minus log(Qtn)]2 + [122 + log(049)]2
Ic = radic[3
Qtn = 27947 n = 041 lt 10 Ic = 14
Layer 3
From Figure 10 fs = 12 psi and qc = 1500 psi
20
qt = qc + u2 ∙ (1 minus a) = 1500 psi + 0 psi ∙ (1 minus 08) = 1500 psi
= 1445 psf + γt ∙ 11 feet = 1445 psf + 1172 pcf ∙ 11 feet = 2734 psf = 190 psi σvo
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 081
(qt minus σvo) (1500 psi minus 190 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = γwater ∙ (z minus 119911119908) = 6224 pcf ∙ (23 feet minus 17 feet) = 3734 psf = 99 psi
prime σvo = σvo minus uo = 190 psi minus 99 psi = 164 psi
(qt minus σvo)σatm (1500 psi minus 190 psi )145 psi = = Qtn prime )n (σvoσatm (164 psi145 psi)n
prime σvo 164 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(081)]2
Qtn = 947 n = 06 lt 10 Ic = 19
21
Layer 4
From Figure 10 fs = 7 psi and qc = 1200 psi
qt = qc + u2 ∙ (1 minus a) = 1200 psi + 0 psi ∙ (1 minus 08) = 1200 psi
= 2734 psf + γt ∙ 11 feet = 2734 psf + 1121 pcf ∙ 7 feet = 3519 psf = 244 psi σvo
fs 7 psi Fr() = 100 ∙ = 100 ∙ = 060
(qt minus σvo) (1200 psi minus 244 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = γwater ∙ (z minus 119911119908) = 6224 pcf ∙ (30 feet minus 17 feet) = 8091 psf = 130 psi
prime σvo = σvo minus uo = 244 psi minus 130 psi = 188 psi
(qt minus σvo)σatm (1200 psi minus 244 psi )145 psi = = Qtn prime )n (σvoσatm (188 psi145 psi)n
prime σvo 188 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(060)]2
22
Qtn = 686 n = 064 lt 10 Ic = 19
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic Qtn and Fr the first layer is defined as a ldquoDrained Gravelly Sandrdquo from
Figure 11
Figure 11 Soil layer 1 using SBT method
23
Layer 2
Based on values of Ic Qtn and Fr the second layer is defined as a ldquoDrained Sandrdquo from Figure
12
Figure 12 Soil layer 2 using SBT method
24
Layer 3
Based on values of Ic Qtn and Fr the third layer is defined as a ldquoDrained Sandrdquo from Figure 13
Figure 13 Soil layer 3 using SBT method
25
Layer 4
Based on values of Ic Qtn and Fr the fourth layer is defined as a ldquoDrained Sandrdquo from Figure 14
Figure 14 Soil layer 4 using SBT method
Effective Stress Friction Angle
Layer 1
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(3533) = 456deg
Layer 2
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(2795) = 445deg
26
Layer 3
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(947) = 393deg
Layer 4
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(686) = 378deg
Stress History
Layer 1
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (13frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(24139 kPa minus 345 kPa)072 = 4717 kPa
= 684 psi
prime σp 684 psi YSR = = = 136
primeσvo 50 psi
Layer 2
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + (14 1 + ( frasl ) frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(34132 kPa minus 689 kPa)072 = 605 kPa
= 877 psi
prime σp 877 psi YSR = = = 87
primeσvo 100 psi
27
Layer 3
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + (19 1 + ( frasl ) frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(10342 kPa minus 131 kPa)072 = 254 kPa
= 368 psi
prime σp 368 psi YSR = = = 22
primeσvo 164 psi
Layer 4
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (19frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(8274 kPa minus 168 kPa)072 = 215 kPa = 312 psi
prime σp 312 psi YSR = = = 17
primeσvo 188 psi
Lateral Stress Coefficient
Layer 1
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(456deg)) ∙ 136sin( 456deg) = 18
Layer 2
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(445deg)) ∙ 87sin( 445deg) = 14
28
Dprime 17480 psi
Eprime = = = 15890 psi 11 11
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(393deg)) ∙ 22sin( 393deg) = 06
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(378deg)) ∙ 17sin( 378deg) = 05
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3501 psi minus 50 psi) = 17480 psi
Eprime 15890 psi
Kprime = = = 8828 psi [3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Layer 3
Layer 4
Ground Stiffness and Soil Moduli
Layer 1
Values of qt and fs need to be in MPa
MR = (146(241 MPa)053 + 1355(012 MPa)14 + 236)244 = 3416 MPa = 49575 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1172 kPa m Vs (mfrasls) = [101 ∙ log(24139 kPa) minus 114]167 ∙ (100 ∙ ) = 2745
24139 kPa s
Vs = 9012 fts
29
γt 1204 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 37 ∙ (9012 ) = 304 ∙ 106psf = 21000 psi s
Layer 2
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3500 psi minus 100 psi) = 17480 psi
Dprime 17480 psi Eprime = = = 15890 psi
11 11
Eprime 15890 psi Kprime = = = 8828 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(241 MPa)053 + 1355(012 MPa)14 + 236)244 = 3415 MPa = 49562 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1172 kPa m Vs (mfrasls) = [101 ∙ log(24133 kPa) minus 114]167 ∙ (100 ∙ ) = 2745
24133 kPa s
Vs = 9012 fts
γt 1204 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
30
2 ft Gmax = ρt ∙ Vs
2 = 37 ∙ (9012 ) = 304 ∙ 106psf = 21000psi s
Layer 3
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (1500 psi minus 190 psi) = 7405 psi
Dprime 7405 psi Eprime = = = 6732 psi
11 11
Eprime 6732 psi Kprime = = = 3740 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(103 MPa)053 + 1355(008 MPa)14 + 236)244 = 1506 MPa = 21854 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 827 kPa m Vs (mfrasls) = [101 ∙ log(10343 kPa) minus 114]167 ∙ (100 ∙ ) = 2611
10343 kPa s
Vs = 8566 fts
γt 1172 pcf slug ρt = = = 36
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 36 ∙ (8566 ) = 27 ∙ 106psf = 19000 psi s
Layer 4
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (1200 psi minus 244 psi) = 5878 psi
31
Dprime 5878 psi Eprime = = = 5343 psi
11 11
Eprime 5343 psi Kprime = = = 2969 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(83 MPa)053 + 1355(005 MPa)14 + 236)244 = 165 MPa = 16901 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 483 kPa m Vs (mfrasls) = [101 ∙ log(8274 kPa) minus 114]167 ∙ (100 ∙ ) = 2243
8274 kPa s
Vs = 7360 fts
γt 1121 pcf slug ρt = = = 35
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 35 ∙ (7360 ) = 389 ∙ 106psf = 13000 psi s
32
2122 Example 2 Direct CPT Methods for Geoparameters on Clay
Several geoparameters need to be determined based on the given CPT data collected for the South
Abutment (Figure 15) The groundwater table (GWT) was measured at 60 feet Determine all the
geoparameters found in Table 5 at depths of 0 feet to 42 feet All sand layers can be assumed ldquodrainedrdquo
ν = 02 All clay layers can assume ν = 049
33
Figure 15 CPT data from Minnesota for example problem 2
34
Table 5 Geoparameters
Symbol Parameter Symbol Parameter
γt Soil total unit weight Eʹ Drained Youngrsquos modulus
Ic CPT material index Kʹ Bulk modulus
SBT Soil behavior type (SBT) MR Resilient modulus
ʹ σp Preconsolidation stress Gmax Small-strain shear modulus
YSR Yield stress ratio su Undrained shear strength
ϕʹ Effective friction angle cv Coefficient of consolidation
Ko Lateral stress coefficient k Hydraulic conductivity
Dʹ Constrained modulus
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
35
Figure 16 Soil layers using ldquorules of thumbrdquo
Unit weight of water γw = 6224 pcf
Layer 1
From Figure 16 fs = 13 psi taken as a representative value of the layer
13 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1179 pcf
145 psi
Layer 2
From Figure 16 fs = 12 psi taken as a representative value of the layer
12 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 3
From Figure 16 fs = 20 psi taken as a representative value of the layer
36
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
Layer 4
From Figure 16 fs = 2 psi taken as a representative value of the layer
2 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1004 pcf
145 psi
CPT Material Index
Layer 1
From Figure 10 fs = 13 psi and qc = 3000 psi
qt = qc + u2 ∙ (1 minus a) = 3000 psi + 0 psi ∙ (1 minus 08) = 3000 psi
σvo = γt ∙ 2 feet = 1179 pcf ∙ 2 feet = 235 psf = 16 psi
fs 13 psi Fr() = 100 ∙ = 100 ∙ = 043
(qt minus σvo) (3000 psi minus 16 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 16 psi minus 0 psi = 16 psi
(qt minus σvo)σatm (3000 psi minus 16 psi )145 psi = = Qtn prime )n (16 psi145 psi)n (σvoσatm
prime σvo 16 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(043)]2
37
Qtn = 4126 n = 032 lt 10 Ic = 121 therefore sand
Layer 2
From Figure 10 fs = 12 psi and qc = 250 psi
qt = qc + u2 ∙ (1 minus a) = 250 psi + 10 psi ∙ (1 minus 08) = 252 psi
σvo = 235 psf + γt ∙ 30 feet = 235 psf + 1172 pcf ∙ 30 feet = 3751 psf = 260 psi
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 53
(q minus σ ) (252 psi minus 260 psi) t vo
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 260 psi minus 0 psi = 260 psi
(qt minus σvo)σatm (250 psi minus 260 psi )145 psi = = Qtn prime (σvoσatm)n (260 psi145 psi)n
38
prime σvo 260 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(53)]2
Qtn = 87 n = 10 le 10 Ic = 32 therefore clay
Layer 3
From Figure 10 fs = 20 psi and qc = 4000 psi
qt = qc + u2 ∙ (1 minus a) = 4000 psi + 8 psi ∙ (1 minus 08) = 40016 psi
σvo = 3751 psf + γt ∙ 2 feet = 3751 psf + 1219 pcf ∙ 2 feet = 3995 psf = 277 psi
fs 16 psi Fr() = 100 ∙ = 100 ∙ = 054
(qt minus σvo) (30016 psi minus 277 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 277 psi minus 0 psi = 277 psi
(qt minus σvo)σatm (30016 minus 277)145 psi = = Qtn prime (σvoσatm)n (277 psi145 psi)n
39
primeσvo 277 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(054)]2
Qtn = 1962 n = 052 le 10 Ic = 15 therefore sand
Layer 4
From Figure 10 fs = 2 psi and qc = 250 psi
qt = qc + u2 ∙ (1 minus a) = 250 psi + 0 psi ∙ (1 minus 08) = 250 psi
= 3994 psf + γt ∙ 8 feet = 3994 psf + 1004 pcf ∙ 8 feet = 4798 psf = 333 psi σvo
fs 2 psi Fr() = 100 ∙ = 100 ∙ = 092
(qt minus σvo) (250 psi minus 333 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 333 psi minus 0 psi = 333 psi
(qt minus σvo)σatm (250 psi minus 333 psi )145 psi = = Qtn prime (σvoσatm)n (333 psi145 psi)n
40
prime σvo 333 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(092)]2
Qtn = 65 n = 10 lt 10 Ic = 29 therefore clayey silt
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic Qtn and Fr the first layer is defined as a ldquoDrained Sandrdquo from Figure 17
41
Figure 17 Soil layer 1 using SBT method
Layer 2
Based on values of Ic Qtn and Fr the second layer is defined as a ldquoUndrained Clayrdquo from Figure
18
42
Figure 18 Soil layer 2 using SBT method
43
Layer 3
Based on values of Ic Qtn and Fr the third layer is defined as a ldquoDrained Sandrdquo from Figure 19
Figure 19 Soil layer 3 using SBT method
44
Layer 4
Based on values of Ic Qtn and Fr the fourth layer is defined as a ldquoUndrained Silty Mixrdquo from
Figure 20
Figure 20 Soil layer 4 using SBT method
Effective Stress Friction Angle
Layer 1 (sand)
45
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(4126) = 464deg
Layer 2 (clay)
qt minus σvo
ϕprime(deg) = 295deg ∙ Bq0121 ∙ [0256 + 0336 ∙ Bq + log ( )]
primeσvo
Where
(u2 minus uo) (10 psi minus 0 psi) Bq = = = 004
(qt minus σvo) (252 psi minus 260 psi)
252 psi minus 260 psi ϕprime(deg) = 295deg ∙ 00440121 ∙ [0256 + 0336 ∙ 0044 + log ( )]
260 psi
ϕprime(deg) = 245deg
Layer 3 (sand)
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(1962) = 428deg
Layer 4 (clay)
qt minus σvo
ϕprime(deg) = 295deg ∙ Bq0121 ∙ [0256 + 0336 ∙ Bq + log ( )]
primeσvo
Where
(u2 minus uo) (5 psi minus 0 psi) Bq = = = 002
(qt minus σvo) (251 psi minus 333 psi)
252 psi minus 333 psi ϕprime(deg) = 295deg ∙ 0020121 ∙ [0256 + 0336 ∙ 002 + log ( )]
333 psi
ϕprime(deg) = 265deg
Stress History
46
Layer 1
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (12frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(3000 psi minus 33)072 = 1051 psi
prime σp 1051 psi YSR = = = 642
primeσvo 16 psi
Layer 2
028 028
prime m = 1 minus = 1 minus = 099 25 25 1 + (
Icfrasl ) 1 + (32frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(252 psi minus 260)099 = 734 psi
prime σp 734 psi YSR = = = 28
primeσvo 260 psi
Layer 3
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + ( frasl ) 1 + (15frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(4002 psi minus 277)072 = 1288 psi
47
prime σp 1288 psi YSR = = = 46
primeσvo 277 psi
Layer 4
028 028
prime m = 1 minus = 1 minus = 097 25 25 Ic 1 + ( frasl ) 1 + (29frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(250 psi minus 333)097 = 626 psi
prime σp 626 psi YSR = = = 19
primeσvo 333 psi
Lateral Stress Coefficient
Layer 1
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(464deg)) ∙ 642sin( 464deg) = 56
Layer 2
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(245deg)) ∙ 28sin( 245deg) = 090
Layer 3
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(428deg)) ∙ 46sin( 428deg) = 091
Layer 4
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(266deg)) ∙ 19sin( 266deg) = 073
48
Ground Stiffness and Soil Moduli
Layer 1
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3000 psi minus 16 psi) = 14991 psi
Dprime 14991 psi Eprime = = = 13629 psi
11 11
Eprime 13629 psi Kprime = = = 7571 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(207 MPa)053 + 1355(009 MPa)14 + 236)244 = 2816 MPa = 40873 psi
fs
03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 896 kPa m Vs (mfrasls) = [101 ∙ log(20685 kPa) minus 114]167 ∙ (100 ∙ ) = 2564
20685 kPa s
Vs = 8412 fts
γt 1179 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
2 ft 2 Gmax = ρt ∙ Vs = 37 ∙ (8412 ) = 260 ∙ 106psf = 17992 psi
s
Layer 2
49
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (252 psi minus 260 psi) = 11298 psi
Dprime 11298 psi Eprime = = = 10271 psi
11 11
Eprime 10271 psi Kprime = = = 17118 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(049))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(17 MPa)053 + 1355(008 MPa)14 + 236)244 = 443 MPa = 64309 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 827 kPa m Vs (mfrasls) = [101 ∙ log(17375 kPa) minus 114]167 ∙ (100 ∙ ) = 2646
17375 kPa s
Vs = 8680 fts
γt 1172 pcf slug ρt = = = 36
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 36 ∙ (8680 ) = 274 ∙ 106psf = 19036 psi s
Layer 3
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (4002 psi minus 277 psi) = 19869 psi
Dprime 19869 psi Eprime = = = 18063 psi
11 11
50
Eprime 18063 psi Kprime = = = 10035 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(276 MPa)053 + 1355(014 MPa)14 + 236)244 = 4017 MPa = 58309 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1379 kPa m Vs (mfrasls) = [101 ∙ log(27591 kPa) minus 1379]167 ∙ (100 ∙ ) = 2854
27591 kPa s
Vs = 9363 fts
γt 121 pcf slug ρt = = = 38
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 38 ∙ (9363 ) = 33 ∙ 106psf = 23050 psi s
Layer 4
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (250 psi minus 333 psi) = 1088 psi
Dprime 1088 psi Eprime = = = 989 psi
11 11
Eprime 989 psi Kprime = = = 16490 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(049))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
51
MR = (146(17 MPa)053 + 1355(001 MPa)14 + 236)244 = 360 MPa = 52189 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 138 kPa m Vs (mfrasls) = [101 ∙ log(17238 kPa) minus 114]167 ∙ (100 ∙ ) = 1545
17238 kPa s
Vs = 5069 fts
γt 1004 pcf slug ρt = = = 31
ga 322 ftfrasls2 ft3
2 ft 2 Gmax = ρt ∙ Vs = 31 ∙ (5069 ) = 80 ∙ 105psf = 5565 psi
s
Undrained Shear Strength
Layer 2
qt minus σv0 250 psi minus 26 psi su = = = 187 psi
12 12
Layer 4
qt minus σv0 250 psi minus 333 psi su = = = 181 psi
12 12
Coefficient of Consolidation
52
Example dissipation data are shown in Figure 21 Example calculations are provided
Figure 21 Dissipation t50 data
Layer 1
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (7197)075
cv = = = 359 cmsec 1 sec t50
Layer 2
53
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (94392)075
cv = = = 0008 cmsec 3000 sec t50
Layer 3
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (8303)075
cv = = = 665 cmsec 06 sec t50
Layer 4
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (26714)075
cv = = = 087 cmsec 11 sec t50
Hydraulic Conductivity
Layer 1
cm 1 psi 359 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 10 ∙ 10minus4 cmsec Dprime 14984 psi
Layer 2
cm 1 psi 0008 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 73 ∙ 10minus7 cmsec Dprime 11379 psi
Layer 3
cm 1 psi 665 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 20 ∙ 10minus5 cmsec Dprime 148650 psi
54
Layer 4
cm 1 psi 087 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 18 ∙ 10minus4 cmsec Dprime 10861 psi
55
CHAPTER 3 DIRECT CPT METHOD FOR SHALLOW
FOUNDATIONS
31 PROCEDURE
Shallow foundation analysis is typically done in a two-part traditional procedure The traditional
techniques are no longer required as a direct CPT method for square rectangular and circular shallow
footings is available (Figure 22) This process has the soil types grouped into four main categories sands
silts fissured clays and intact clays When determining soil types for each design it is believed footings
on sands and silts act in a fully drained manner while intact clays act in an undrained manner under
conditions of constant volume In order to determine the vertical stress-displacement-capacity of
square rectangular and circular shallow footings follow the steps provided towards the solution given
by Equation 27
Figure 22 Direct CPT method for shallow foundations
Equation 27 may be used to calculate all footing stresses from zero to the bearing capacity (qmax) To
calculate qmax for a sized footing of width (B) length (L) and thickness (t) follow steps 1 through 7
provided The settlement (s) can be determined after the calculation of qmax by simply rearranging
56
Equation 27 where the allowable stress (qallow) is defined as qmax divided by the factor of safety (FS) For
shallow footings a FS value of 3 is commonly used in geotechnical engineering
120782120787 minus120782120785120786120787 119956 119923 119954119950119938119961 = 119945119956 ∙ 119954119957119951119942119957 ∙ (119913
) ∙ (119913
) 27 119950119938119961
2 119902119898119886119909 0345 1 frasl119865119878 119871 119904 = 119861 ∙ [ ∙ ∙ ( ) ] 28
119861 ℎ119904 119902119905119899119890119905
Where hs = the foundation soil formation parameter qtnet = the net corrected cone tip resistance
311 Step 1 Estimating Footing Dimensions
In Minnesota frost heave can have devastating effect on a shallow foundations It is common practice to
place a foundation bearing elevation below the expected maximum frost depth (roughly 45 to 6 feet
below ground level) With this assumption estimate a footing size (B x L) for design to obtain
representative data from CPT roughly 15middotB below the foundation depth (Df) as shown in Figure 23 The
CPT data collected will consist of
cone tip resistance (qc)
measured porewater pressure acting behind the cone tip (u2) as shown in Figure 24
sleeve friction (fs)
57
Figure 23 Direct CPT method introduction
Figure 24 Differentiation of porewater pressure measurement locations (Lunne et al 1997)
312 Step 2 Soil Characterization
The soil behavior type (SBT) and CPT material index (Ic) govern the value of the formation factor hs The
first step is following the steps in Soil Unit Weight to determine soil layering and γt values for each
layer To determine a representative unit weight (γsoil) for all the layers in the range of Df to 15∙B below
58
Df the user will need to use their engineering judgement on the definition of ldquorepresentative unit
weightrdquo This single value of unit weight is used in further calculations such as the total vertical soil stress (σvo) from Equation 29 and effective vertical stress (σvo) from Equation 30 Values of σvo and σvo
will need to be calculated at 15middotB below Df
120648119959119952 = sum(120632119956119952119946119949 ∙ 119963) 29
prime 120648119959119952 = 120648119959119952 minus 119958119952 30
Where z = Df +15middotB uo=γwatermiddot(z-zw)
The qc will need to be corrected using Equation 31 in the case of fine-grained soils that develop excess porewater pressure during cone penetration These values will be used in the following calculations
119954119957 = 119954119940 + 119958120784 ∙ (120783 minus 119938) 31
Where 119860119899 a = cone area ratio = eg MnDOT commonly uses a = 08 119860119888
119860119899 = cross-sectional area of load cell or shaft 119860119888 = projected area of the cone
The cone area ratio is determined based on the type of piezocone tip used during in-situ field testing
Manufacturer specifications should provide the measured net area ratio (a) for the particular cone
penetrometer as determined by calibration in a pressurized triaxial chamber
313 Step 2a Foundation soil formation parameter
With the representative value γsoil continue to follow the steps in CPT Material Index through Soil
Behavior Type (SBT) to better define the type of soil at depth 15∙B below Df Use the value of Ic
calculated at depth 15∙B below Df to calculate hs with Equation 36 This parameter is based on the soil
type with typical values shown in Figure 25 The data on silts and sands are considered fully drained
whereas the fissured clay subset may be partially drained to undrained
59
23 ℎ119904 = 28 minus
119868119888 15 32
1+( ) 24
Figure 25 Foundation soil formation parameter hs versus CPT material index Ic (Mayne 2017)
314 Step 3 Soil elastic modulus and Poissonrsquos ratio
Details about the soil such as its elastic modulus (Es) and Poissonrsquos ratio (ν) will be needed for further
calculations A representative value for the Es can be determined from in-situ field tests Values of ν can
be assigned as 02 for drained sands and as 05 for undrained loading cases involving clays (Jardine et al
1985 Burland 1989)
315 Step 4 Net cone tip resistance
Calculate qtnet the mean value of net cone tip resistance 15middotB below the foundation bearing elevation
using Equation 33
33 119902119905119899119890119905 = 119902119905 minus 120590119907119900
60
316 Step 5 Bearing capacity of the soil
Use the assumed B and L calculated hs and qtnet to determine the soils bearing capacity (qmax) from
Equation 34 Use Table 6 to calculate qmax by using the maximum allowable settlement ratio (sB)max
correlating to the soil type If hs is in between the given values interpolate to acquire (sB)max
Table 6 Bearing capacity defined by soil type
Type of Soil hs (sB)max
Clean Sands 058 12
Silts 112 10
Fissured Clays 147 7
Intact Clays 270 4
05 minus0345 119904 119871 119902119898119886119909 = ℎ119904 ∙ 119902119905119899119890119905 ∙ (119861
) ∙ (119861
) 34 119898119886119909
317 Step 6 Settlement
Settlement can be calculated directly using the results from Equation 35 A FS equal to 3 is common in
foundation engineering
2 119902119898119886119909 0345 1 frasl119865119878 119871 119904 = 119861 ∙ [ ∙ ∙ ( ) ] 35
119861 ℎ119904 119902119905119899119890119905
318 Step 7 Final Check
Check that the applied stress (q) is less than qmax using Equation 36 Repeat the process again with a
new B andor new L if q gt qmax
05 119904 ) 119902 = ℎ119904 ∙ 119902119905119899119890119905 ∙ (119861
minus0345 119871 ∙ ( )
11986136
61
32 EXAMPLE PROBLEMS
321 Example 3 Direct CPT Method on Sands
A footing size needs to be determined based on the given CPT data collected for the South Abutment
(Figure 27) The footing stress (q) was determined to be 8000 psf Estimate a footing size (B x L) and
determine the bearing capacity of the foundation (qmax) using the direct CPT method provided Also
determine the expected settlement based on the calculated bearing capacity
Figure 26 Diagram of footing profiles for Example 3
The soil elastic modulus determined from seismic CPT (SCPT) are shown in Table 7
Table 7 SCPT Results
Depth (feet) Es (tsf)
Bottom of layer
3 557
6 433
9 557
12 695
17 590
22 501
27 571
32 505
62
Figure 27 CPT data from Northern Minnesota
Solution
Step 1 Assume the footing is placed below the frost depth of 6 feet
Estimate L L = 50 feet = 600 inches
63
Estimate B B = 12 feet = 144 inches
Estimate footing thickness t t = 2 feet
Df = 6 feet = 72 inches
Df + 15middotB = 24 feet = 288 inches
Step 2 Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 28 Schematic of foundation design associated with CPT data
Layer 1
From Figure 28 fs = 20 psi taken as a representative value of the layer
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
64
Layer 2
From Figure 28 fs = 20 psi taken as a representative value of the layer
20psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
Layer 3
From Figure 28 fs = 8 psi taken as a representative value of the layer
8 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1134 pcf
145 psi
Using the unit weights calculated between Df and 15B determine a representative unit weight of the soil to calculate the total and effective soil stresses Based on layer 3 being the weakest supporting layer γsoil = 113 pcf
σvo = γsoil ∙ (119863119891 + 15 ∙ B) = 113 lbsfraslft3 ∙ 24 feet = 2721 psf = 189 psi
prime σvo = σvo minus uo = 2721 psf minus 0 psf = 2721 psf = 189 psi
Calculate the cone tip resistance between Df and Df + 15B below the foundation depth
qt = qc + u2 ∙ (1 minus a) = 1250 psi + 0 psi ∙ (1 minus 08) = 1000 psi
Step 2a CPT Material Index
Using Figure 28 the sleeve friction at 15B below the foundation depth is about 16 psi
65
fs 16 psi Fr() = 100 ∙ = 100 ∙ = 13
(qt minus σvo) (1250 psi minus 189 psi)
Iterate to solve for Ic Steps are not shown for brevity
(qt minus σvo)σatm (1250 psi minus 189 psi )145 psi = = Qtn prime (σvoσatm)n (189 psi145 psi)n
prime σvo 189 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Qtn = 703 n = 072 lt 10 Ic = 210
66
Figure 29 Soil type for example problem 3
Use the Ic value to determine the foundation soil formation parameter
Calculate the foundation soil formation parameter The tip resistance is about 1250 psi at 24 feet and
the cone area ratio was determined to be 08 from information provided by the manufacturer
23 23 hs = 28 minus = 28 minus = 078 = SandSilt 15 15 Ic 210
1 + ( ) 1 + ( ) 24 24
67
Step 3 Determine representative values of soil elastic modulus from soil testing and Poissons ratio The
soil elastic modulus was taken as the average value between depths of Df and 15middotB (6 feet and 24 feet)
433 + 557 + 695 + 590 + 501 Es = = 555 tsf = 708 psi
5
ldquoDrained sandsiltrdquo gives a ν = 020
Step 4 Calculate the net cone tip resistance
qtnet = qt minus σvo = 1250 psi minus 189 psi = 12311 psi
Step 5 Calculate the bearing capacity of the sand
In drained sandssilts the bearing capacity is taken as the stress when (sB) = 011 (or 11 foundation width)
minus0345 minus0345 05 s L 600 qmax = hs ∙ qtnet ∙ ( ) ∙ ( ) = 078 ∙ (9795) ∙ (011)05 ∙ ( )
B B 144
= 193 psi = 27860 psf qmax
Assuming a factor of safety (FS) of 3
qmax = 645 psi = 9287 psf
3
Step 6 Calculate settlement
68
119902119898119886119909 0345 2
1 frasl119865119878 L 119904 = 119861 ∙ [ ∙ ∙ ( ) ]
ℎ119904 119902119905119899119890119905 B
2 0345 1 645 psi 600 inches s = 144 inches ∙ [ ∙ ∙ ( ) ] = 18 inches
078 12311 psi 144 inches
Step 7 Determine if q gt qmax
q = 8000 psf lt 9287psf = qmax3
69
CHAPTER 4 DIRECT CPT METHOD FOR DEEP FOUNDATIONS
41 INTRODUCTION
The axial compression capacity (Qtotal) for a single pile includes a side component (Qside) end bearing
component (Qbase) and pile weight (Wpile) as shown in Figure 30 Since piles commonly push through
several layers a summation of the unit side frictions acting on the pile segments must be considered
over the length of the pile While the examples shown in this Guide resemble hand calculations
computer software is more efficient Programming the procedure is possible and represents a practical
method of designing deep foundations However commercial software is frequently available and is
MnDOTrsquos most common method of designing deep foundations
There are upwards of 40 different direct CPT methods that have been developed over the past five
decades to determine a piles axial compression capacity Many of the earliest methods relied on hand-
recorded information where qc data from mechanical CPTs would be collected at 20 cm intervals
whereas the direct CPT method uses scaled penetrometer readings via specified algorithms to obtain
the pile unit side friction and unit end bearing The method that will be used for deep foundation design
is the Modified UniCone method which uses all three readings of the electronic piezocone (qt fs and u2)
while addressing a variety of pile foundation types
Figure 30 Direct CPT evaluation of axial pile capacity
70
The modified UniCone Method is based upon a total of 330 pile load tests (three times the original
Unicone database) that were associated with SCPTu data Originally the UniCone method provided
approximate soil classification in five groups via a chart of qE vs fs (Figure 31) where qE = qt -u2 Later
using the modified approach with a larger data set provided soil sub classifications as shown in Figure
32 This new 9-zone normalized soil behavior type is determined using CPT data in combination with the
CPT Material Index
Figure 31 UniCone Method soil behavior type using CPT (Mayne 2017)
Figure 32 Modified UniCone Method soil behavior type using CPT (Mayne 2017)
71
42 MODIFIED UNICONE METHOD
In order to determine the axial pile capacity using the modified method the first step requires the
determination of geoparameters as shown in Direct CPT Method for Soil Characterization Once the
soil unit weight and CPT Material index are determined for each soil layer then the effective cone
resistance pile unit side friction (fp) and pile end bearing resistance (qb) can be determined
421 Step 1
Work through the steps provided in CPT Method for Soil Characterization until each soil layer is
defined by its CPT material index and soil behavior type using Figure 6 After these steps have been
completed continue to Step 2 to determine qE qb and fp
422 Step 2
Once qt is determined for each layer the effective cone resistance can be calculated using Equation 37
119902119864 = 119902119905 minus 1199062 37
Where (a) qE is the specific value at each elevation along the pile sides for determining fp and (b) at the bottom of the pile qE is averaged in the vicinity of the pile tip from the tip bearing elevation to about one diameter beneath the tip for determining qb
Using CPT material index and qE the pile end bearing resistance is calculated using Equation 38
38 119902119887 = 119902119864 ∙ 10(0325∙119868119888minus1218)
The pile unit side friction is obtained from qE and Ic
39 119891119901 = 119902119864 ∙ 120579119875119879 ∙ 120579119879119862 ∙ 120579119877119860119879119864 ∙ 10(0732∙119868119888minus3605)
Where θPT = coefficient for pile type (084 for bored 102 for jacked 113 for driven piles) θTC = coefficient for loading direction (111 for compression and 085 for tension) θRATE = rate coefficient applied to soils in SBT zone 1 through 7 (109 for constant rate of penetration test and 097 for maintained load tests)
72
423 Step 3
Due to piles extending through multiple layers the unit side components acting on various pile
segments would need to be summed if not using the direct CPT method Since CPT calculates data at
regular intervals of 2 cm to 5 cm along the sides of the pile the average fp in each layer can be used
directly in Equation 40 to obtain the shaft capacity
119876119904119894119889119890 = 119891119901 ∙ 119860119904 40
Where fp = average pile side friction along pile length from eqn 39 As = πmiddotdmiddotH d = pile diameter and H = length embedded below grade
The base capacity for a pile in compression loading is given by Equation 41 For piles in tension (or uplift)
Qbase can be taken as 0
= 119902119887 ∙ 119860119887 41 119876119887119886119904119890 Where
qb = end bearing resistance from eqn 38 Ab = πmiddotd24 (area of a circular pile)
424 Step 4
The final step is to calculate the axial pile capacity using Equation 42
119876119905119900119905119886119897 = 119876119904119894119889119890 + 119876119887119886119904119890 minus 119882119901119894119897119890 42
43 AXIAL PILE DISPLACEMENTS
Movement of pile foundations can be assessed using elastic continuum theory which has been
developed using finite element analyses boundary elements and analytical closed-form solutions In
the case of piles passing through several soil layers the elastic solution can be used by stacking pile
segments (each with its own stiffness) as represented by soil Youngrsquos modulus The use of software is
recommended for pile groups Several available programs such as DEFPIG GROUP and PIGLET can
handle pile groups under axial and lateralmoment loading
73
44 EXAMPLE PROBLEMS
441 Example 5 Direct CPT Methods Axial Pile Capacity
Several piles need to be placed beneath the edge of a building Use the CPT data collected for this site
shown in Figure 34 to determine the axial capacity for one of the piles Assume round steel driven piles
will be used with a diameter of 1275 inches wall thickness of 025 inches and lengths of 80 feet
Concrete will be used to fill the piles To solve for all geoparameters use the Direct CPT Method for Soil
Characterization section All sand layers can be assumed ldquodrainedrdquo with ν = 02
Figure 33 Deep foundation end bearing pile diagram
74
Figure 34 CPT data from Minnesota for example problem 5
75
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 35 Soil layers using ldquorules of thumbrdquo for pile capacity example using direct CPT method
Layer 1
From Figure 35 fs = 18 psi taken as a representative value of the layer
76
18 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1210 pcf
145 psi
Layer 2
From Figure 35 fs = 12 psi taken as a representative value of the layer
12psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 3
From Figure 35 fs = 20 psi taken as a representative value of the layer
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
CPT Material Index
Layer 1
From Figure 35 fs = 18 psi and qc = 3000 psi
qt = qc + u2 ∙ (1 minus a) = 3000 psi + 20 psi ∙ (1 minus 08) = 3004 psi
∙ 4 feet = 1219 pcf ∙ 4 feet = 4877 psf = 34 psi σvo = γt
fs 18 psi Fr() = 100 ∙ = 100 ∙ = 060
(qt minus σvo) (3004 psi minus 34 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
77
uo = 0 psi
prime σvo = σvo minus uo = 34 psi minus 0 psi = 34 psi
(qt minus σvo)σatm (3004 psi minus 34 psi )145 psi = = Qtn prime )n (σvoσatm (34 psi145 psi)n
prime σvo 34 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(060)]2
Qtn = 3593 n = 038 lt 10 Ic = 14 (ie sand)
Layer 2
From Figure 35 fs = 12 psi and qc = 500 psi
qt = qc + u2 ∙ (1 minus a) = 500 psi + 40 psi ∙ (1 minus 08) = 508 psi
= 4838 psf + γt ∙ 45 feet = 4838 psf + 1172 pcf ∙ 45 feet = 5756 psf = 400 psi σvo
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 260
(qt minus σvo) (508 psi minus 400 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
78
uo = 0 psi
prime σvo = σvo minus uo = 400 psi minus 0 psi = 400 psi
(qt minus σvo)σatm (508 psi minus 400 psi )145 psi = = Qtn prime (σvoσatm)n (400 psi145 psi)n
prime σvo 400 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(260)]2
Qtn = 117 n = 10 le 10 Ic = 29 (ie clayey silt)
Layer 3
From Figure 35 fs = 20 psi and qc = 5000 psi
qt = qc + u2 ∙ (1 minus a) = 5000 psi + 0 psi ∙ (1 minus 08) = 5000 psi
= 5756 psf + γt ∙ 6 feet = 5756 psf + 1219 pcf ∙ 6 feet = 6488 psf = 451 psi σvo
fs 20 psi Fr() = 100 ∙ = 100 ∙ = 040
(qt minus σvo) (5000 psi minus 451 psi)
79
Step 2 Iterate to solve for Ic Steps are not shown for brevity
prime σvo = σvo minus uo = 451 psi minus 0 psi = 451 psi
(qt minus σvo)σatm (5000 psi minus 451 psi )145 psi = = Qtn prime )n (σvoσatm (451 psi145 psi)n
prime σvo 451 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(040)]2
= 1801 n = 06 lt 10 = 15 (ie sand) Qtn Ic
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic (14) Qtn (359) and Fr (06) the first layer is defined as a ldquoGravelly Sandrdquo
from Figure 36
80
Figure 36 Soil layer 1 using SBT method
81
Layer 2
Based on values of Ic (29) Qtn (117) and Fr (26) the second layer is defined as a ldquoSilty Mixrdquo
from Figure 37
Figure 37 Soil layer 2 using SBT method
82
Layer 3
Based on values of Ic (15) Qtn (180) and Fr (04) the third layer is defined as a ldquoSandrdquo from
Figure 38
Figure 38 Soil layer 3 using SBT method
Effective Cone Resistance
Layer 1
83
qE = qt minus u2 = 3004 psi minus 20 psi = 2984 psi
Layer 2
qE = qt minus u2 = 508 psi minus 40 psi = 468 psi
Layer 3
qE = qt minus u2 = 5000 psi minus 0 psi = 5000 psi
Pile End Bearing Resistance
Layer 3
= qE ∙ 10(0325∙Icminus1218) qb = 5000 psi ∙ 10(0325∙15minus1218) = 9085 psi
Pile Unit Side Friction
Layer 1
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 2984 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙14minus3605) = 99 psi
Layer 2
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 468 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙29minus3605) = 211 psi
Layer 3
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 5000 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙15minus3605) = 202 psi
84
Axial Pile Capacity
The side capacity is based on pile design as shown in Figure 39
Layer 1
Qside = fp ∙ As = 99 psi ∙ (π ∙ d ∙ L) = 99 psi ∙ (π ∙ 1275 in ∙ 48 in) = 19064 lb
Layer 2
Qside = fp ∙ As = 211 psi ∙ (π ∙ d ∙ L) = 211 psi ∙ (π ∙ 1275 in ∙ 588 in) = 457122 lb
Layer 3
Qside = fp ∙ As = 202 psi ∙ (π ∙ d ∙ L) = 202 psi ∙ (π ∙ 1275 in ∙ 240 in) = 58170 lb
Figure 39 Soil layering compared to pilings
85
End Bearing Resistance in Layer 3
d2 1275 in2
Qbase = qb ∙ Ab = 9085 psi ∙ (π ∙ ) = 9085 psi ∙ (π ∙ ) = 115932 lb 4 4
Wp = (γsteel ∙ Asteel ∙ Depth) + (γsteel ∙ Aconc ∙ Depth)
Wp = (490 pcf ∙ 003 ft2 ∙ 55 ft) + (150 pcf ∙ 085 ft2 ∙ 55 ft) = 7724 lbs
Layer 3
Qtotal = ΣQside + Qblayer3 minus Wp
Qtotal = (19064 + 45713 + 58170) lbs + 115932 lbs minus 7724 lbs = 642563 lbs
= 642 kips Qtotal
Table 8 Summary of selected and calculated parameters
Layer L (in) qc
(psi)
fs
(psi)
Fr Ic Qtn qE
(psi)
qb
(psi)
Qbase
(lb)
fp
(psi)
Qside
(lb)
Qtotal
(kip)
1 48 3000 18 06 14 356 2984 NA NA 99 19064
2 588 500 12 26 29 117 468 NA NA 211 457122
3 240 5000 20 04 15 180 5000 9085 115932 202 58170
Total 115932 534355 642
86
REFERENCES
Abu-Farsakh MY Zhang Z Tumay M and Morvant M (2008) Computerized cone penetration test for soil classification Transportation Research Record 2053 47-64
Agaiby SA (2018) Advancements in the interpretation of seismic piezocone tests in clays and other geomaterials PhD dissertation School of Civil amp Environmental Engineering Georgia Institute of Technology Atlanta GA
Agaiby SS and Mayne PW (2018) Interpretation of piezocone penetration and dissipation tests in sensitive Leda Clay at Gloucester Test Site Canadian Geotechnical Journal (accepted for publication 08 March 2018) CGJ-2017-0388
Amar S Baguelin F Canepa Y and Frank R (1998) New design rules for the bearing capacity of shallow foundations based on Menard PMT Geotechnical Site Characterization Vol 2 (Proc ISC-1 Atlanta) 727-733
American Petroleum Institute (API) (1981) Planning designing and constructing fixed offshore platforms Recommended practice API-RP2A 17th edition Washington DC API
American Petroleum Institute (API) (1987) Planning designing and constructing fixed offshore platforms Recommended practice API-RP2A 18th edition Washington DC API
American Petroleum Institute (API) (1989) Planning designing and constructing fixed offshore platforms Recommended practice API-RP2A 19th edition Washington DC API
Andersen KH and Stenhamar P (1982) Static plate loading tests on overconsolidated clay Journal of the Geotechnical Engineering Division ASCE 108 (GT7) 919-934
Basu D and Salgado R (2012) Load and resistance factor design of drilled shafts in sands Journal of Geotechnical amp Geoenvironmental Engineering 138 (12) 1455-1469
Berardi R Jamiolkowski M and Lancellotta R (1991) Settlement of shallow foundations in sands selection of stiffness on the basis of penetration resistance Geotechnical Engineering Congress Vol 1 185-200 (Proc GSP-27 Boulder) Reston Virginia ASCE
Brand EW Muktabhant C and Taechathummarak A (1972) Load tests on small foundations in soft clay Performance of Earth and Earth-Supported Structures Vol 1 Part 2 903-928 (Proc PC) ASCE Reston Virginia ASCE
Briaud JL and Gibbens RM (1999) Behavior of five large spread footings in sand Journal of Geotechnical amp Geoenvironmental Engineering 125 (9) 787-797
Briaud JL (2007) Spread footings in sand load-settlement curve approach Journal of Geotechnical amp Geoenvironmental Engineering 133 (8) 905-920
Brown DA Turner JP and Castelli RJ (2010) Drilled Shafts Construction Procedures and LRFD Design Methods Geotechnical Engineering Circular GEC 10 Report No FHWA NHI-10-016 Washington DC National Highway Institute US DOT
87
Burland J B (1973) Shaft Friction of Piles in Clay -A Simple Fundamental Approach Ground Eng 6(3) 30-42
Cai G Liu S and Puppala A (2012) Reliability assessment of CPTu-based pile capacity predictions in soft clay deposits Engineering Geology 84-91
Canadian Geotechnical Society (1992) Canadian Foundation Engineering Manual Third Edition Vancouver British Columbia BiTech Publishers
Cerato AB and Lutenegger AJ (2007) Scale effects of shallow foundation bearing capacity on granular material Journal of Geotechnical amp Geoenvironmental Engineering 133 (10) 1192-1201
Chen BS-Y and Mayne PW (1996) Statistical relationships between piezocone measurements and stress history of clays Canadian Geotechnical Journal 33 (3) 488-498
Cho W and Finno RJ (2010_ Stress-strain responses of block samples of compressible Chicago glacial clays Journal of Geotechnical amp Geoenvironmental Engineering 136 (1) 178-188
Chung SF Randolph MF and Schneider JA (2006) Effect of penetration rate on penetrometer resistance in clay Journal of Geotechnical amp Geoenvironmental Engineering 132 (9) 1188-1196
Clausen CJF Aas PM and Karlsrud K (2005) Bearing capacity of driven piles in sand the NGI approach Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis Group
Consoli NC Schnaid F and Milititsky J (1998) Interpretation of plate load tests on residual soil site Journal of Geotecnical amp Geoenvironmental Engingeering 124 (9) 857-867
Dasenbrock DD (2006) Assessment of pile capacity by static and dynamic methods to reconcile design predictions with observed performance Proc 54th Minnesota Geotechnical Engineering Conference 95-108 edited by Labuz J Univ of Minnesota St Paul
DeBeer E and Martens A (1957) Method of computation of an upper limit for the influence of heterogeneity of sand layers on the settlement of bridges 275-282 Proc ICSMFE-4 Vol 1 London ICSMFE
DeBeer EE (1965) Bearing capacity and settlement of shallow foundations on sand Proc Symposium on Bearing Capacity and Settlement of Foundations 15-33 Duke University Durham NC
Decourt L (1999) Behavior of foundations under working load conditions Proc PCSMGE-11 Foz do Iguassu Brazil Vol 4 453-487
Demers D and Leroueil S (2002) Evaluation of preconsolidation pressure and the overconsolidation ratio from piezocone tests of clay deposits in Quebec Canadian Geotechnical Journal 39 (1) 174-192
Eslaamizaad S and Robertson PK (1996) Cone penetration test to evaluate bearing capacity of foundations in sand 429-438 Proc CGCFG-49 Vol 1 St Johns Newfoundland
Eslami A Alimirzaei M Aflaki E and Molaabasi H (2017) Deltaic soil behavior classification using CPTu records Marine Georesources amp Geotechnology 35 (1) 62-79
88
Eslami A and Gholami M (2005) Bearing capacity analysis of shallow foundations from CPT data 1463-1466 Proc ICSMGE- 16 Vol 3 (Osaka) Millpress Rotterdam
Eslami A and Gholami M (2006) Analytical model for the ultimate bearing capacity of foundations from cone resistance Scientia Iranica 13 (3) 223-233
Eslami A and Fellenius BH (1997) Pile capacity by direct CPT and CPTu methods applied to 102 case histories Canadian Geotechnical Journal 34 (6) 886-904
Fellenius BH (2009) Basics of Foundation Design Vancouver BiTech Publishers
Fellenius BH and Altaee A (1994) Stress and settlement of footings in sand Vertical and Horizontal Deformations of Foundations and Embankments 2 1760-1773
Fellenius BH and Eslami A (2000) Soil profile interpreted from CPTu data Proc Geotechnical Engineering Conference Year 2000 Geotechnics Asian Inst of Technology Bangkok Thailand Available from wwwfelleniusnet
Finno R J and Chung C K (1992) Stress-strain-strength responses of compressible Chicago glacial clays Journal of Geotechnical Engineering 118 (10) 1607ndash1625
Fleming K Weltman A Randolph M and Elson K (2009) Piling Engineering Third Edition London Taylor amp Francis Group
Frank R and Magnan J-P (1995) Cone penetration testing in France national report Proceedings International Symposium on Cone Penetration Testing 3 (CPT95 Linkoumlping) Swedish Geotechnical Society Report 3(95) 147-156
Gavin K Adekunte A and OKelly B (2009) A field investigation of vertical footing response on sand Geotechnical Engineering 162 257-267
Hegazy YA (1998) Delineating geostratigraphy by cluster analysis of piezocone data PhD dissertation School of Civil amp Environmental Engineering Georgia Institute of Technology Atlanta GA
Holtz RD Kovacs WD and Sheehan TC (2011) An Introduction to Geotechnical Engineering 2nd Edition London Pearson Publishing
Hunt CE Pestana JM Bray JD and Riemer M (2002) Effect of pile driving on static and dynamic properties of soft clay Journal of Geotechnical amp Geoenvironmental Engineering 128 (1) 13-24
Jamiolkowski M Ladd CC Germaine JT and Lancellotta R (1985) New developments in field and lab testing of soils Proc ICSMFE-11 1 57-154
Jardine R Chow F Overy R and Standing J (2005) ICP design methods for driven piles in sands and clays London Thomas Telford Publishing
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89
Karlsrud K (2012) Prediction of load-displacement behavior and capacity of axially-loaded piles in clay based on analyses and interpretation of pile load test results PhD Dissertation Norwegian Univ Science amp Technology Trondheim Norway
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Karlsrud K Lunne T Kort DA and Strandvik S (2001) CPTU correlations for clays Proc 16th ICSMGE 2 683-702
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Knappett JA and Craig RF (2012) Craigs Soil Mechanics 8th Edition London Spon Press Taylor amp Francis
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Kulhawy FH and Mayne PW (1990) Manual for Estimating Soil Properties for Foundation Design EPRI Report EL-6800 Electric Power Research Institute Palo Alto 306 page Download from wwwepricom
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Ku T and Mayne PW (2015) In-situ lateral stress coefficient (K0) from shear wave velocity measurements in soils Journal of Geotechnical amp Geoenvironmental Engineering 141 (12) 101061(ASCE) GT1943-56060001354
Ladd CC and DeGroot DJ (2003) Recommended practice for soft ground site characterization Soil amp Rock America 2003 3-57
Larsson R (1997) Investigations and Load Tests in Silty Soils SGI Report R-54 Linkoumlping Swedish Geotechnical Institute
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90
Lehane BM (2003) Vertically loaded shallow foundation on soft clayey silt Geotechnical Engineering 156 (1) Thomas Telford London 17-26
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Mayne PW (1991) Determination of OCR in Clays by Piezocone Tests Using Cavity Expansion and Critical State Concepts Soils and Foundations 31 (2) 65-76
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91
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92
Mlynarek Z Wierzbicki J Goglik S and Bogucki M (2014) Shear strength and deformation parameters of peat and gyttja from CPTu DMT and VST Proceedings 5th International Workshop on CPTu and DMT in soft clays and organic soils 193-210 Poznan Poland Polish Committee on Geotechnics Menard Polska Tensar Internationa
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Nowacki F Karlsrud K and Sparrevik P (1996) Comparison of recent tests on OC clay and implications for design Large Scale Pile Tests in Clay London Thomas Telford
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Poulos HG and Davis E (1980) Pile Foundation Analysis amp Design New York John Wiley amp Sons
Powell JJM and Lunne T (2005) Use of CPTU data in clays and fine-grained soils Studia Geotechnica et Mechanica XXVII(3) 29-66
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93
Robertson PK (2009) Cone penetration testing a unified approach Canadian Geotechnical J 46 (11) 1337-1355
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Robertson PK and Cabal KL (2015) Guide to Cone Penetration Testing for Geotechnical Engineering 6th Edition Signal Hill California Gregg Drilling amp Testing Inc
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Schmertmann JH (1978) Guidelines for cone penetration test performance and design Report FHWA-TS-78-209 Washington DC Federal Highway Administration
Schnaid F Wood WR Smith AKC and Jubb P (1993) Investigation of bearing capacity and settlements of soft clay deposits at Shellhaven Predictive Soil Mechanics London Thomas Telford
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Senneset K Sandven R amp Janbu N (1989) Evaluation of soil parameters from piezocone tests In-Situ Testing of Soil Properties for Transportation Transportation Research Record 1235 24-37
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Tomlinson MJ (1986) Foundation Design amp Construction London Longman Scientific
Tomlinson MJ (1987) Pile Design and Construction Practice London Viewpoint Publication
Trofimenkov JG (1974) Penetration testing in the USSR Proceedings ESOPT-1 1 147-154
Tumay MT Hatipkarasulu Y Marx ER and Cotton B (2013) CPTPCPT-based organic material profiling Proc 18th Intl Conf on Soil Mechanics amp Geotechnical Engineering Paris 633-636
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Uzielli M and Mayne PW (2012) Load-displacement uncertainty of vertically-loaded shallow footings on sands and effects on probabilistic settlement estimation Georisk Assessment and Management of Risk for Engineered Systems and GeoHazards 6 (1) 50-69
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Van Dijk BFJ and Kolk HJ (2011) CPT-based design method for axial capacity of offshore piles in clay Frontiers in Offshore Geotechnics II 555-560
Vesić AS (1975) Bearing capacity of shallow foundations Foundation Engineering Handbook New York Van Nostrand Reinhold Publishers
Vesic AS (1977) NCHRP Synthesis of Highway Practice 42 Design of Pile Foundations Washington DC Transportation Research Board
Yu F and Yang J (2012) Base capacity of open-ended steel pipe piles in sand J Geotechnical and Geoenvironmental Engineering 138 (9) 1116-1128
Zawrzykraj P Rydelek P and Bakowska A (2017) Geoengineering properties of Eemian peats from Radsymin (central Poland) in the light of static cone penetration and dilatometer tests Engineering Geology 226 290-300
95
APPENDIX A
List of Figures
Figure A1 Conventional methods versus direct CPT approach to shallow foundation response A-3
Figure A2 Direct bearing capacity relationship for embedded square and strip footings situated on clean
sands (after Schmertmann 1978) A-6
Figure A3 Direct CPT bearing factor Rk as function of footing width B and embedment depth from finite
element analyses (Tand et al 1995) A-6
Figure A4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio and sand
consistency (Eslaamizaad amp Robertson 1996) A-7
Figure A5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp Salgado (2005) in
terms of footing size relative density and base settlement A-8
Figure A6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and normalized
cone tip resistance (after Eslami amp Gholami 2005) A-8
Figure A7 Direct relationship between ultimate bearing stress in clays and measured cone tip resistance
(Schmertmann 1978) A-10
Figure A8 Direct relationship between ultimate foundation bearing stress in clays and net cone tip
resistance (after Tand et al 1986) A-11
Figure A9 Measured load-displacements for each of the five large footings at TAMU (Briaud amp Gibbens
1994) sand site A-12
Figure A10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU (Briaud amp
Gibbens 1994) footings A-13
Figure A11 Characteristic stress versus square root normalized displacements for TAMU footings A-15
Figure A12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sand A-16
Figure A13 Summary of sand formation factors with corresponding CPT resistances A-19
Figure A14 Normalized foundation stresses vs square root of normalized displacement for spread
footings on sand (modified after Mayne et al 2012) A-20
Figure A15 Definitions of capacity from three types of observed foundation load test responses
(modified after Kulhawy 2004) A-21
Figure A16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footings A-22
A-1
Figure A17 Measured load vs displacement curves for 3 shallow foundations on clay at Baytown Texas
(Stuedlein amp Holtz 2010) A-25
Figure A18 Characteristic stress vs square root of normalized displacement for all three foundations at
Baytown site (Stuedlein amp Holtz 2010) A-26
Figure A19 Normalized foundation stress vs square root of normalized displacements A-27
Figure A20 Applied foundation stress vs CPT-calculated stress for 67 footings A-28
Figure A21 Applied footing stress vs CPT calculated stress on logarithmic scales A-29
Figure A22 Summary graph and equations of direct CPT method for evaluating the vertical A-30
Figure A23 Foundation soil formation parameter hs versus CPT material index Ic A-32
Figure A24 Bearing capacity ratio qmaxqnet versus CPT material index Ic A-33
Figure A25 Influence factor for rectangular foundations from elastic theory solution (Mayne amp
Dasenbrock 2017) A-34
List of Tables
Table A1 Direct CPT methods for bearing capacity of footings on clean sands A-4
Table A2 Direct CPT methods for bearing capacity of footings on clays A-9
Table A3 Summary of large footings soil conditions and reference sources of database A-17
Table A4 Sources for shallow foundation performace A-34
A-2
A1 INTERPRETATION OF SOIL PARAMETERS FROM CONE
PENETRATION TESTS
A11 Introduction
Geotechnical site characterization is an important first step towards the evaluation of
subsurface conditions and determination of soil layering geomaterial classification and the
evaluation of soil engineering parameters for the analysis and design of foundations retaining
walls tunnels excavations embankments and slope stability Increasing use of the electronic
cone penetrometer in highway applications is prominent in the USA because the results are
obtained much faster and less expensive than traditional methods that rely on rotary drilling
augering sampling and laboratory testing Moreover the cone penetration test (CPT) provides
at least three independent and continuous readings on soil behavior that are digitally recorded
and fully available at the end of the sounding As such the reliable and consistent
interpretation of CPT data is important since civil engineering works will best realize the
efficiency and economy of this technology in constructed facilities for county state and
interstate projects
A111 Cone Penetration Testing
The cone penetrometer is an electronically-instrumented steel probe that is vertically pushed into the
ground at constant rate of 20 mms (08 insec) using a hydraulic system The penetrometer is outfitted
with load cells pressure transducers and inclinometers that measure at least three readings
continuously with depth (a) cone tip resistance qt (b) sleeve friction fs and (c) porewater pressure u2
With the latter reading the device is called a piezocone thus the designation CPTu is used to indicate
porewater pressure Additional sensors are available that can be used to measure inclination
resistivity pH temperature shear wave velocity and other readings if desired
Figure A1 shows a schematic rendering of the cone penetration test that is performed in the
field per ASTM D 5778 procedures The hydraulic system is nominally rated at 180 kN (20 tons)
capacity and often mounted on a truck but can also be positioned on tracked vehicles or
anchored frames Figure A2 shows a MnDOT cone truck that uses the full dead weight of the
rig for the hydraulic reaction forces which are applied at the center of the vehicle The cab is
enclosed so that soundings may proceed during inclement weather and the data acquisition
system and operator are protected
Figure A 1 Setup and procedures for co ne penetration testing (CPT)
A-4
Figure A2 CPT truck-mounted rig used by MnDOT
A representative sounding taken at the I-35 test site just northeast of Saint Paul is presented in Figure
A3 Here the separate profiles of qt fs and u2 with depth are shown in side-by-side plots In the last
graph the hydrostatic porewater pressure (uo) due to the groundwater table is indicated by the blue
dashed line
These readings are used to interpret the soil layers types and properties of the ground as
discussed in the following sections
A-5
Figure A3 Representative CPT sounding at I-35 test site northeast of St Paul MN showing profiles of
(a) cone resistance (b) sleeve friction and (c) penetration porewater pressures
A112 Soil Parameter Interpretation
A variety of soil engineering parameters can be interpreted from CPT results on the basis of
theoretical frameworks analytical models or numerical simulations otherwise by empirical
methods based on correlations and statistics of databases Figure A4 shows a conceptual
utilization of the readings from the CPT for interpretation of geostratigraphy compressibility
flow characteristics and stress-strain-strength behavior of soils
Table A1 lists a selection of geoparameters that have been addressed for these purposes The
most common or useful relationships will be discussed in subsequent sections and reference is
made to other sources (Lunne et al 1997 Mayne 2007 Robertson amp Cabal 2015) for additional
details
A-6
Figure A4 Conceptual post-processing of CPT results for geoparameter evaluation
Table A1 List of common geoparameters interpreted from CPT data
Symbol Parameter Remarks Notes
SBT Soil behavior type (SBT)
Ic CPT material index
γt Unit weight
ρt Mass Density
σvo Overburden stress
u0 Hydrostatic pressure
A-7
σvo Effective stress
Table A1 Continued
Symbol Parameter Remarks Notes
σp Preconsolidation stress σp = 033 (qnet)m where m = fctn (Ic)
(or Effective Yield Stress) where stresses are in kPa
YSR Yield stress ratio YSR = σpσvo (formerly OCR)
c Effective cohesion intercept
ϕ Effective friction angle
su Undrained shear strength
D Constrained modulus
G0 Small-strain modulus
G Shear modulus
E Youngs modulus
cvh Coefficient of consolidation
K Bulk modulus
A-8
LI
K0 Lateral stress coefficient
k Hydraulic conductivity
Liquefaction index
A12 Geostratigraphy and Soil Behavioral Type (SBT)
In routine CPT soil samples are not normally taken and thus the measured stress friction andor
porewater pressure readings are used to infer the types of soils This can be accomplished using
three basic approaches
a Rules of Thumb or approximate guidelines for a quick visual assessment
b Soil Behavioral Type (SBT) Charts that are based on either the three readings (qt fs u2) net readings
including net cone resistance (qnet = (qt-σvo) effective cone resistance (qE = qt - u2) friction ratio (Rf =
100middotfsqt) and excess porewater pressures or using normalized CPT readings such as Q F Bq or U
c Probabilistic methods where the CPT readings have been calibrated from lab tests on
recovered soil samples (eg Tumay et al 2013)
A121 Approximate Rules of Thumb
The approximate rules of thumb provide simple guidelines for soil type and suggest that sands are
identified when qt gt 5 MPa and u2 asymp u0 whereas intact clays are prevalent when qt lt 5 MPa and u2 gt u0
(Mayne et al 2002) The magnitude of porewater pressures reflects the consistency of the intact clay
such that soft (u2 asymp 2middotu0) firm (u2 asymp 4middotu0) stiff (u2 asymp 8middotu0) and hard (u2 asymp 20middotu0) However for fissured
overconsolidated clays the measured porewater pressures are less than hydrostatic in fact often
negative u2 lt 0 On land negative porewater pressure readings at the shoulder filter element of the
piezocone should be greater than -1 bar ie u2 gt -100 kPa (-147 psi) Also for clean sands the friction
ratio (FR = 100middotfsqt lt 1) while for insensitive clays (FR gt 4)
A-9
Figure A5 presents a 36-m deep piezocone sounding from the MnDOT Wakota Bridge site which crosses
the Mississippi River at I-494 (Dasenbrock 2006) The qt and u2 readings have been annotated using the
aforementioned rules of thumb indicating the predominance of sands at this site As noted there are
five interbedded clay layers evident at depths of 1 m 3-4 m 7 - 12 m 17 m and 22-27 m
Figure A5 Interpretation of soil types from CPT data taken from the Wakota Bridge on I-494 using
approximate rules of thumb
A122 Soil Behavioral Type Charts
The most common approach for identification of soil types is based on soil behavioral charts
Reviews of several chart methods are provided elsewhere (Kulhawy amp Mayne 1990 Fellenius amp
Eslami 2000 Shahri et al 2015) Current popularity favors the 9-zone classification system to
evaluate soil behavioral type (SBT) that uses normalized piezocone readings (Robertson 1990
1991 Lunne et al 1997)
A-10
(119902119905 minus 120590119907119900) (a) normalized tip resistance 119876 = A11 prime 120590119907119900
100∙119891119904 (b) normalized sleeve friction 119865() = A12
(119902119905 minus 120590119907119900)
(1199062 minus 1199060) (c) normalized porewater pressure 119861119902 = A13
(119902119905 minus 120590119907119900)
While the Q-F-Bq classification is a three-dimensional plot it is often presented as a set of two-dimensional
plots specifically Q vs F and Q vs Bq as shown in Figure A6 Data are grouped according to layers and
can be superimposed to identify their association with the nine soil behavioral types All indirect CPT soil
classification approaches should be cross-checked and verified for a particular geologic setting and local
geotechnical conditions before routine use in practice
Figure A6 Nine-zone soil behavioral type charts (a) normalized cone resistance vs normalized
sleeve friction (b) normalized cone resistance vs normalized porewater pressure parameters
(adapted after Robertson 1991 Lunne et al 1997)
The development of a CPT material index (Ic) has been found advantageous in the initial screening of soil
types and helps to organize the sounding into their respective SBT zones of similar soil response In this
case the CPT index is found from (Robertson 2009)
A-11
A2 22 )log221()log473( rtnc FQI
where Qtn = modfied stress-normalized net cone resistance given by
(119902119905 minus 120590119907119900)120590119886119905119898 119876 = A31 prime (120590119907119900120590119886119905119898)119899
where σatm = reference stress equal to atmospheric pressure (1 atm asymp 1 bar = 100 kPa asymp 1 tsf asymp 147 psi)
The exponent n is soil-type dependent n = 1 (clays) n asymp 075 (silts) and n asymp 05 (sands) If units of bars are
used then Qtn (units of bars) is simply
(119902119905 minus 120590119907119900) 119876 = 119899 A32
prime ) (120590119907119900
The operational value of exponent n is found by iteration Initially an exponent n = 1 is used to calculate
the starting value of Ic (ie Qtn = Q) and then the exponent is upgraded to
prime 120590119907119900 119899 = 0381 ∙ 119868119888 + 005 ( ) minus 015 le 10 A4 120590119886119905119898
Then the index Ic is recalculated Iteration converges quickly and generally only 3 cycles are needed to
secure the operational Ic at each depth The soil zones and associated Ic values are detailed in Figure A7
The sensitive soils of zone 1 can be screened using the following expression
A-12
Zone 1 119876119905119899 lt 12119890119909119901( minus14 ∙ 119865119903)
A5
If any soil layers are found within zone 1 (sensitive soils) then caution should be exercised as these
structured clays are prone to instability collapses and difficulties in construction and performance
Another indicator of zone 1 sensitive soils is when the normalized porewater parameter Bq gt 08 When
these criteria are evident the geoengineer should consult with senior geotechnical personnel or the chief
engineer for guidance
Figure A7 Delineation of soil behavioral type zones using CPT material index Ic
A-13
Stiff overconsolidated clayey sands and sandy clays soils of zone 8 (15 lt Fr lt 45) and zone 9 (Fr gt
45) can be identified from the following criterion
Zones 8 and 9 0020)1(00030)1(0050
12
rr
tnFF
Q A6
The remaining soil types are identified by the CPT material index Zone 2 (organic clayey soils Ic ge 360)
Zone 3 (clays to silty clays 295 le Ic lt 360) Zone 4 (silt mixtures 260 le Ic lt 295) Zone 5 (sand mixtures
205 le Ic lt 260) Zone 6 (clean sands 131 le Ic lt 205) and Zone 7 (gravelly to dense sands Ic le 131)
The red dashed line at Ic = 260 represents an approximate boundary separating drained (Ic lt 260) from
undrained behavior (Ic gt 260)
Once the specific zone has been assigned to a layer a visual representation can be made to show
either the zone number or colorization so that the predominant layers and soil types can be
realized
A123 Probabilistic Soil Types from CPT
The inference of soil type has been addressed using probabilistic methods as discussed by Abu-
Farshakh et al (2008) and Tumay et al (2013) Here the CPT readings can be post-processed to
provide the percentage of sand silt and clay components with depth The output is consistent
and compatible with the current MnDOT soil classification chart (Figure A8) which is similar in
content to that used by the USDA
A-14
Figure A8 Chart for soil classification system by MnDOT
A13 Identifying Soft Organic Soils
The identification of soil type using cone penetration tests (CPT) is usually done on the basis of empirical
soil behavioral type (SBT) charts that use the measured readings (qt fs andor u2) While there at least
25 different sets of charts available (ie Hegazy 1998 Eslami et al 2017 Valsson 2016) one of the most
widely used and popular is the 9-zone SBT system that uses three normalized cone readings Qtn Fr and
Bq (Robertson amp Cabal 2007 Lunne et al 1997) The system is denoted as SBTn and plotted on charts in
terms of Q vs F and Q vs Bq as shown in Figure A9
A-15
1
10
100
1000
01 1 10
Norm
aliz
ed
Con
e T
ip R
esis
tan
ce
Q
Normalized Friction Fr = 100 fst(qt - svo) ()
9-Zone Soil Behavioral Type Chart
Gravelly Sands(zone 7)
Sands
(zone 6)
Sandy Mixtures(zone 5)
Silt Mix(zone 4)
Clays
(zone 3)
Organic Soils(zone 2)
Sensitive Clays
and Silts(zone 1)
Very stiff OC clayto silt (zone 9)
Very stiffOC sandto clayey
sand(zone 8)
1
10
100
1000
-06 -04 -02 00 02 04 06 08 10 12 14
No
rmal
ized
Co
ne
Tip
Res
ista
nce
Q
Normalized Porewater Bq = Du2(qt-svo)
Clays(zone 3)
Sensitive(zone 1)Organic
(zone 2)
Gravelly Sands(zone 7)
Sands(zone 6)
Figure A9 CPT charts for SBTn system (a) Q versus F (b) Q versus Bq
Of particular concern in geotechnical site characterization is the proper identification of soft
organic soils such as organic clays and silts (OL and OH) and peats (Pt) These soils often cause
difficulties and problems during construction and performance of civil engineering works because
of bio-degradation long-term creep excessive settlements andor other issues
The SBTn system has a category labelled Zone 2 recognizes organic soils However several recent
studies have noted that data from CPTu soundings do not always fall within the bounds
delineated using Zone 2 even though laboratory tests and field inspections clearly note the
presence of organic soils Lab tests to classify organic soils includes loss on ignition (LOI gt 5)
and two pairs of Atterberg limits testings per ASTM standards as well as the visual-manual
identification of strong organic odor and smell and dark color notably black and dark gray
Results by Mlynarek et al (2014) show that organic soils in Poland are not adequately identified by the
Q-F chart as seen in Figure A10 Moreover studies by Zawrzykraj et al (2017) found that neither the Q-
A-16
F and Q-Bq plots worked well to recognize organic soils and peats in Poland Nejaim et al (2016) found
that Brazilian soft organic clays were misclassified as Zone 3 (clays) rather than Zone 2 (organic clays)
Similarly a recent study on CPTu data from organic clays located at 24 sites in the USA Sweden Mexico
Brazil and Australia have been compiled (Agaiby 2018) These data are shown to generally avoid falling
with the Zone 2 boundaries in either Q-F or Q-Bq charts thus are not recognized during these soundings
as indicated by Figure A11 The exception in this case is the Mexico City clay which is correctly identified
however the other 23 sites are generally not properly recognized and categorized Additional findings by
Missiaen et al (2015) found that the CPTu results in Belgian soils also did not identify organic clays and
peats in the non-normalized version of these charts that uses Qt versus friction ratio (Rf = 100middotfsqt) as
detailed by Robertson amp Cabal (2007)
Figure A10 Soil behavioral chart with superimposed CPTu data from Polish organic soils
(from Mlynarek et al 2014)
A-17
Figure A11 Paired SBTn charts with CPTu data from 24 organic clay sites indicating non-compliance
with Zone 2 (organic soils)
(compiled by Agaiby 2018)
A14 Simplified Yield Stress Evaluation in Clays by CPTu
From the development of a hybrid formulation of spherical cavity expansion (SCE) theory with critical-
state soil mechanics (CSSM) the effective preconsolidation stress or yield stress (σp) of clays can be
expressed in terms of three piezocone parameters (a) net cone tip resistance qnet = qt - σvo (b)
measured excess porewater pressure Δu2 = u2 - u0 and (c) effective cone resistance qE = qt - u2 Details
on the SCE-CSSM solutions are given elsewhere (Mayne 1991 Mayne 2017 Agaiby amp Mayne 2018)
A-18
For normal and well-behaved clays which include inorganic fine-grained soils such as clays and silts
of low sensitivity a set of simple relationships can be derived By adopting characteristic default values
for effective friction angle ϕ = 30deg and rigidity index (IR = 100) linear expressions for the effective
preconsolidation stress are obtained
prime 120590119901 = A7 033 ∙ 119902119899119890119905
prime 120590119901 = 053 ∙ ∆1199062 A8
prime 120590119901 = 060 ∙ 119902119864 A9
A141 Application to soft Chicago clay
An example of a common geomaterial in this category includes the infamous soft Chicago clays
that were deposited in a glacio-lacustrine environment (Cho amp Finno 2010) The soft clay has a
mean water content of 20 liquid limit of 38 and plasticity index PI= 12 Results of CPTu tests
conducted at the national geotechnical experimentation site (NGES) at Northwestern University
are shown in Figure A12 (Ouyang amp Mayne 2017) As seen in the profile a thin soft clay resides
between depths of 9 to 105 m with a thicker soft clay layer in the range of 12 to 18 m
A-19
0
2
4
6
8
10
12
14
16
18
20
00 02 04 06 08 10 12D
ep
th (
m)
CPTu qt and u2 (kPa)
sand (SP)
Blodgett
Park Ridge
Deerfieldsoft clay
ClayLayerof Study
Northwestern UnivNational Test Site
Figure A12 Results of CPTu sounding at the NGES at Northwestern University Illinois
Post-processing of the CPTu data using Equations A7 A8 and A9 show good agreement in evaluating the
profile of effective yield stress in this clay layer compared with results from one-dimensional
consolidation tests made on undisturbed samples taken at the site
A-20
10
11
12
13
14
15
16
17
18
19
20
0 100 200 300 400 500 600 700 800 900 1000
De
pth
(m
)Effective Yield Stress sp (kPa)
Consols
033 qnet
053 Du2
060 qE
svo
Figure A13 Evaluation of effective yield stresses at NWU using consolidation tests and CPTu
A142 Application to soft Bay Mud California
Another example of a well-behaved fine-grained soil is the San Francisco Bay Mud which is a soft clay
deposit in northern California Results of a representative CPTu are shown in Figure A14 and indicate the
soil profile at a site in eastern San Francisco (Hunt et al 2002) For this site index values include 70 lt
wn lt 90 60 lt LL lt 80 and 35 lt PI lt 45 Post-processing the CPTu data using Equations A7 A8 and
A9 show good agreement with each other as well as with the results of consolidation tests as seen in
A-21
Figure A15 The consolidation tests were performed using a CRS device as reported by Pestana et al
(2002)
San Francisco Bay Mud (Pestana et al 2002)
0
5
10
15
20
25
30
35
40
0 1000 2000 3000
De
pth
(m
ete
rs)
Cone Resistance qt (kPa)
0
5
10
15
20
25
30
35
40
0 10 20 30 40 50 60
Sleeve Friction fs (kPa)
0
5
10
15
20
25
30
35
40
0 500 1000 1500
Porewater u2 (kPa)
Miscellaneous Fill
Young Bay Mud
Young Bay Mud
Clayey Sand
Miscellaneous Fill
Young Bay Mud
Young Bay Mud
Figure A14 Representative CPTu in San Francisco Bay Mud (data from Hunt et al 2002)
A-22
San Francisco Bay Mud (Pestana et al 2002)
0
5
10
15
20
25
30
0 100 200 300 400 500 600
De
pth
(m
ete
rs)
Preconsolidation Stress sp (kPa)
svo
033 qnet
053 Du2
060 qE
CRS
svo
Figure A15 Evaluation of effective yield stresses in Bay Mud from consolidation tests and CPTu
A15 CPTu Screening of Soft Organic Clays
For soft organic clays the Equations A7 A8 and A8 will not apply thus can serve as a means to
screen such geomaterials from the soil profile Two examples of CPTu soundings in soft organic
clays are presented to illustrate the approach Additional case records and documentation of
CPTu data in organic clays are given in Agaiby (2018)
A151 Soft organic alluvial soils in Washington DC
Results of in-situ tests including CPTu soundings in soft organic clayey silts along the Potomac River at the
Anacostia Naval Air Station are presented by Mayne (1987) Here the soft soils are categorized as organic
clayey silts (OH) per the Unified Soils Classification Systems (USCS) with mean values (and plus and minus
A-23
one standard deviations) of wn = 68 plusmn 16 LL = 83 plusmn 25 and PI = 37 plusmn 17 A representative
piezocone sounding is depicted in Figure A16 For the post-processing of CPTu data the use of the Q-F
and Q-Bq graphs do not find any points within the Zone 2 category for organic soils When the data are
evaluated to determine the profile of effective yield stress the Equations A7 A8 and A9 do not agree as
shown by Figure A17
0
5
10
15
20
25
0 2000 4000 6000
Dep
th (
met
ers)
Cone Resistance qt (kPa)
0
5
10
15
20
25
0 20 40 60 80 100
Sleeve Friction fs (kPa)
0
5
10
15
20
25
0 200 400 600
Pore Pressure u2 (kPa)
Figure A16 Representative CPTu in soft organic clay along the Potomac River DC
A-24
0
5
10
15
20
25
0 3 6 9D
epth
(m
eter
s)Soil Behaviour Type Zone
Q and Fchart
0
5
10
15
20
25
0 100 200 300 400 500 600
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
033qnet
053 Du2
06 qE
Oed
Figure A17 Post-processing of CPTu data from Anacostia-Bolling (a) SBTn zones using CPT material
index Ic (b) CPTu screening equations for yield stress
Note that the hierarchy of the results from Equations A7 A8 and A9 show that the CPTu
evaluations for preconsolidation stress in soft organic clays indicates
[A10] σp = 053 Δu2 lt 033 qnet lt 060 qE
A152 Soft organic peaty clay in Saint Paul MN
Results of CPTu tests were collected during a training exercise underneath I-35E just northeast
of Saint Paul MN in 2007 All three MnDOT CPT rigs available at the time were used to obtain
in-situ test data at the site The site is well known as having soft organic soils and samples were
obtained from three borings made at the site (Boring ID T523 T542 and T518) Boring logs
A-25
indicate the presence of highly organic silt loams with marl and spongy organic clayey silts
From 12 samples the natural water contents ranged from 30 to 191 with a mean of 112 plusmn
58
A representative piezocone sounding from the site (MnDOT CPTu ID F22Y0703C) is used for
this illustration and is presented as Figure A18 Post-processing these data using the SBTn
algorithms and CPT material index show that the soils classify primarily as Zone 3 (clays) and
Zone 4 (silty mix) thus do not identify the geomaterials properly as Zone 2 (organic soils)
0
5
10
15
0 500 1000 1500 2000
Dep
th (
met
ers)
Cone Tip Resistance
qt (kPa)
0
5
10
15
0 10 20 30 40 50 60
Sleeve Friction
fs (kPa)
0
5
10
15
-100 0 100 200 300 400
Porewater Pressure
u2 (kPa)
Figure A18 MnDOT CPTu sounding at the I-35E test site near St Paul Minnesota
A-26
1
10
100
1000
01 1 10
Norm
aliz
ed
Con
e T
ip R
esis
tan
ce
Q
Normalized Friction Fr = 100 fst(qt - svo) ()
9-Zone Soil Behavioral Type Chart
St PaulGravelly Sands(zone 7)
Sands
(zone 6)
Sandy Mixtures(zone 5)
Silt Mix(zone 4)
Clays
(zone 3)
Organic Soils(zone 2)
Sensitive Clays
and Silts(zone 1)
Very stiff OC clayto silt (zone 9)
Very stiffOC sandto clayey
sand(zone 8)
1
10
100
1000
-06 -04 -02 00 02 04 06 08 10 12 14
No
rmal
ized
Co
ne
Tip
Res
ista
nce
Q
Normalized Porewater Bq = Du2(qt-svo)
Clays(zone 3)
Sensitive(zone 1)Organic
(zone 2)
Gravelly Sands(zone 7)
Sands(zone 6)
Figure A19 Post-processing St Paul sounding using the 9-zone SBTn classification system
Using the recommended approach Figure A20 shows the CPT-evaluated yield stress profiles from
equations A7 A8 and A9 Here again the three estimates of σp do not agree but show the same
hierarchy as detailed in Equation A10
A-27
0
5
10
15
20
0 50 100 150 200 250 300 350 400
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
06 qeff svo
033qnet Consols
054 Du2
Figure A20 CPTu screening of soft organic soils at St Paul test site
A16 CPTu Evaluations of Yield Stress in Soft Organic Soils
Once the presence of soft organic clays and silts is identified using the aforementioned CPT
screening algorithms the evaluation of effective yield stress is conducted using the procedure
outlined in Chapter 2 of this MnDOT manual For soft organic soils an exponent of m = 090 is
recommended (Mayne et al 2009) Therefore the preconsolidation stress is obtained from
(units of kPa)
prime = A11 120590119901 033 ∙ (119902119899119890119905)090
The algorithm can be expressed in dimensionless form by
prime = A12 120590119901 033 ∙ (119902119905 minus 120590119907119900)119898prime ∙ (120590119886119905119898100)1minus119898prime
A-28
where σatm = reference stress (= 100 kPa = 147 psi) and m = 090 for soft organic silts and clays
The approach is applied to the two former example case studies in Figure A21 (Washington DC)
and Figure A22 (St Paul MN) respectively showing good agreement with benchmark results
taken from laboratory consolidation tests on undisturbed samples of these sites in both cases
0
5
10
15
20
25
0 100 200 300 400 500 600
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
033qnet^090
Oed
Figure A21 Profiles of yield stress from consolidation tests and CPTu method for organic clayey silts at
Anacostia-Bolling site in Washington DC
A-29
0
5
10
15
20
0 20 40 60 80 100 120 140 160
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
CPT
Consols
120590119901prime = 033 119902119899119890119905
090 in kPa
Figure A22 Profiles of yield stress from consolidation tests and CPTu method for organic clayey silts at
MnDOT test site near Saint Paul Minnesota
A17 Soil Unit Weight
As shown by Figure A23 total soil unit weight (t) can be estimated from the CPT sleeve friction resistance
(Mayne 2014) The equation can be expressed by
120574119905 = 120574119908 ∙ [122 + 015 ∙ 119897119899 (100 lowast 119891119904120590119886119905119898 + 001)] A13
where w = unit weight of water and satm = atmospheric pressure
A-30
Figure A23 Evaluation of soil unit weight from CPT sleeve friction
A18 Effective Stress Friction Angle
Sands
The effective stress friction angle () is one of the most important soil properties as it governs the strength
of geomaterials as well as affects soil-pile interface and pile side friction While an effective cohesion
intercept (c) can also be considered this is usually reserved for cemented or bonded geomaterials or
unsaturated soils and may lose its magnitude with time ageing or with prolonged environmental
changes
For clean quartz to silica sands where porewater pressures are essentially hydrostatic (Bq = 0) the
following expression has been calibrated with triaxial compression test results from undisturbed sand
samples and normalized cone resistances as presented in Figure A24 (Mayne 2007 2014)
A-31
120601prime(119889119890119892) = 176deg + 110deg 119897119900119892 (1199021199051) A14
Figure A24 Evaluation of effective stress friction angle in quartz-silica sands from CPT
Note Relationship applies to drained soil behavior when Ic lt 26 andor Bq lt 01
where qt1 is an earlier form of stress normalized cone tip resistance for sands given by
(119902119905 120590119886119905119898) 1199021199051 = A151 05 prime (120590119907119900120590119886119905119898)
A-32
In terms of units of bars this is expressed simply as (in units of bars)
119902119905 1199021199051 = 05 A152 prime ) (120590119907119900
Recently Robertson amp Cabal (2015) recommended the use of the modified form of normalized cone
resistance (Qtn) in Equation A10
120601prime (119889119890119892) = 176deg + 110deg 119897119900119892 (119876119905119899) A16
In sands qnet asymp qt since the overburden term is small relative to the cone tip resistance Figure A25 shows
that the two normalizations give very comparable results
Figure A25 Comparable values from qt1 and Qtn normalization schemes for CPTs in sands
Clays and Silts
A-33
In the case of soft to firm intact clays and silty clays the effective friction angles are determined from the
normalized cone resistance and porewater pressure parameters (Senneset et al 1989 Mayne 2016) as
shown in Figure A26 The exact solution when the angle of plastification = 0 is given as
qBQ
)tan1(tan61
1)tanexp()245(tan 2
A17
Approximate NTH Solution for from CPTu
qBQ
)tan1(tan61
1)tanexp()245(tan2
Approx asymp 295deg∙Bq0121∙[0256 + 0336∙Bq + log Q ]
1
10
100
20 25 30 35 40 45
Co
ne
Res
ista
nce
Nu
mb
er Q
(degrees)
Bq
010204060810
c = 0 = 0deg
Theory (dots)
Approximation (lines)
Theory
Figure A26 Evaluation of effective stress friction angle in clays and silts from CPTu results
Note Relationship applies to undrained soil behavior when Ic ge 26 andor Bq ge 01
which can be approximately inverted into the form (Mayne 2007)
A-34
0121 120601prime = 295deg ∙ 119861119902 ∙ [ 0256 + 0336 ∙ 119861119902 + log(119876)] A18
This algorithm is specifically applicable for the following ranges of porewater pressure parameter (01 le
Bq lt 1) and effective stress friction angles (20deg le lt 45deg)
A19 Stress History
The stress history can be characterized by an apparent yield stress or preconsolidation stress (sp) as well
as by its normalized and dimensionless form YSR = spsvo = yield stress ratio The YSR is in effect the
same as overconsolidation ratio (OCR) however now generalized to accommodate mechanisms of
preconsolidation that occur beyond just erosion glaciation and removal of overburden stresses but also
due to ageing desiccation repeated cycles of wetting-drying bonding repeated freeze-thaw cycles
groundwater changes and other factors
A-35
A generalized approach for evaluating the yield stress or preconsolidation in soils using net cone
resistance has been formulated as presented in Figure A27 (Mayne 2015) Yield stress in soils from CPT
10
100
1000
10000
10 100 1000 10000 100000
Yie
ld S
tre
ss
s
y
(kP
a)
Net Cone Resistance qt - svo (kPa)
Blessington
General Trend
sy = 033(qt-svo)m
Fissured Clays m = 11
Intact clays m = 10 Sensitive clays m = 09
Silt Mixtures m = 085Silty Sands m = 080
Clean Sands m = 072
m = 11 10 09
085
080
072
Units kPa
Fissured Clays
Figure A27 Evaluation of yield stress or preconsolidation stress in soils from CPT
The algorithm can be expressed in dimensionless form by
1minus119898prime prime 120590119886119905119898 120590119901 = 033 ∙ (119902119905 minus 120590119907119900)119898prime
( ) A19 100
where m = exponent depends on soil type m = 1 (intact clays) 085 (silts) 080 (sandy silts to silty sands)
and 072 (sands) For fissured clays the exponent m may be 11 or higher depending upon the age of the
A-36
formation and degree of jointing and discontinuities Note that fissured clays can be identified when Ic lt
26 and Bq lt 01
The exponent m has also been calibrated with CPT material index as presented in Figure A28
An algorithm to express this relationship for non-fissured soils is given by
25)652(1
2801
cIm
A20
This allows the post-processing of CPT to automatically choose the appropriate exponent m for
a layer by layer analysis
Figure A28 Yield stress exponent m in terms of CPT material index Ic
A-37
A110 Lateral Stress Coefficient
The horizontal geostatic state of stress is represented by the lateral stress coefficient K0 = shosvo
commonly referred to as the at-rest condition The magnitude of K0 for soils that have been loaded and
unloaded can be approximately estimated from
119870119900 = (1 minus 119904119894119899120601prime) ∙ 119874119862119877119904119894119899120601prime A21
Data compiled from in-situ K0 measurements using self-boring pressuremeter tests (SBPMT) total stress
cells (TSC) or push-in spade cells andor laboratory methods (instrumented consolidometers triaxials
andor suction measurements) on a variety of clays silts and sands have been compiled and reported by
Ku amp Mayne (2015) as shown in Figure A29 verifying Equation A15 as a means for evaluating K0 in soils
Figure A29 Relationship between lateral stress coefficient K0 and YSR or OCR in soils (Ku amp Mayne
2015)
A-38
A111 Undrained Shear Strength
The loading of soils can occur under conditions of being fully drained (Du = 0) partially drained or full
undrained (DVV0) where Du = excess porewater pressures (above hydrostatic) and DVV0 = volumetric
strain Further details on specific stress paths are best explained in terms of critical state soil mechanics
or CSSM (Mayne et al 2009 Holtz Kovacs amp Sheahan 2011) The prevailing drainage conditions depend
upon the rate of loading and permeability characteristics of the soil Normally in sands that are pervious
and exhibit high permeability a drained response occurs Exception to this may occur in loose sands
during fast earthquake loading resulting in soil liquefaction In clays that exhibit low permeability a fast
rate of loading will result in undrained loading at constant volume This in fact is a temporary and
transient condition often termed short term loading In the long term eventually porewater pressures
will dissipate (albeit slowly) and a drained response will prevail thus termed long term loading
The overall soil strength is controlled by the effective stress strength envelope Most commonly this is
represented by a simple linear relationship termed the Mohr-Coulomb criterion where the maximum
shear stress (max called the shear strength) is given as
= 119888prime + 120590prime ∙ 119905119886119899 120601prime A22 120591119898119886119909
where c = effective cohesion intercept s = effective normal stress and = effective stress friction angle
As a starting point values of c = 0 and = 30deg can be adopted for all soil types at least until the specific
soil type geologic formation and results from high-quality laboratory or field test data are available
Peak Undrained Shear Strength from Stress History
Using simplified critical state soil mechanics the peak undrained shear strength (max = su) can be
evaluated from the effective stress strength envelope (c = 0 effective friction angle ) and stress history
(ie OCR) in the form
prime DSS 119904119906 = (119904119894119899120601prime2) ∙ (119874119862119877)Λ ∙ 120590119907119900 A23
A-39
which corresponds to a simple shear mode Figure A30 presents the aforementioned relationship
together with data from 17 different clays tested in simple shear
Figure A30 Normalized undrained shear strength from simple shear tests on various clays
versus YSR or OCR
For the triaxial compression (TC) mode the equation would give slightly higher strengths that are
calculated from
prime TC 119904119906 = (1198721198882) ∙ (1198741198621198772)Λ ∙ 120590119907119900 A24
where Mc = (6 middot sin)(3-sin)
Peak Undrained Shear Strength from CPTu
A more direct approach to assessing undrained shear strength is via bearing capacity theory
whereby
A-40
A25 kt
votu
N
qs
s
where Nkt = bearing factor that depends on mode of shearing sensitivity of the clay degree of
overconsolidation and other factors For soft offshore clays the back figured Nkt from data collected at
14 well-documented sites (Low et al 2010) determined mean values based on mode of shearing Nkt =
119 (triaxial compression) Nkt = 136 (simple shear) and Nkt = 133 (vane)
A recent study of 51 clays that were tested by both field piezocone and laboratory CAUC triaxial tests
showed that essentially Nkt = 12 for intact clays and clayey silts of low to medium sensitivity (Mayne
Peuchen amp Baltoukas 2015) Figure A31 shows the slightly different trends for offshore versus onshore
clays For sensitive clays a lower value of Nkt = 10 would be appropriate and for fissured clays a higher
value of Nkt would be in the range of 20 to 30
A-41
A-42
Figure A31 Database from 51 clays with CPT qnet versus lab measured triaxial shear strength
(Mayne Peuchen amp Baltoukas 2015)
For soft to firm clays an alternate means to evaluate undrained shear strength is via the excess porewater
pressures ( u = u2 - u0) from the following
u
uN
us
D
D 2
A26
where N u = porewater bearing factor also dependent on the aforementioned factors The study by
Low et al (2010) found N u = 59 (triaxial compression) N u = 69 (simple shear) and N u = 71 (vane
shear)
A third alternate is found from the effective cone resistance (qE = qt - u2) whereby
kE
tu
N
uqs 2
A27
where NkE is a bearing term for this approach Mayne et al (2015) indicated a mean value of NkE = 8 for
soft-firm intact clays
Remolded Undrained Shear Strength from CPT
The remolded undrained shear strength (sur) is obtained from either field vane tests lab mini-vane shear
tests or lab fall cone devices This affords the evaluation of the clay sensitivity (St) which is defined as the
ratio of peak to remolded strengths at a given water content
119904119906(119901119890119886119896) 119878119905 = frasl A28 119904119906119903
For the CPT the sleeve friction has been noted to give values that are comparable to the
remolded undrained shear strength (Powell amp Lunne 2015) Therefore
asymp 119891119904 A29 119904119906119903
A112 Ground Stiffness and Soil Moduli
The stiffness of the ground can be represented by several geoparameters including the compressibility
indices (Cr Cc Cs) spring constants (kv) and moduli Regarding the latter there are several moduli that
are used in geotechnical engineering including shear modulus (G and Gu) Youngs modulus (E and Eu)
constrained modulus (D) bulk modulus (K) resilient modulus (MR) and subgrade reaction modulus (ks)
A-43
Soil Modulus
The definition of modulus is taken as E = DsD with Figure A32 showing a typical deviator stress (q = s1
- s3) versus axial strain (a) curve Theoretical interrelationships between the elastic moduli G E D and
K have a dependence on the Poissons ratio v The value of Poissons ratio can be taken as vu = 05 for
undrained loading (ie constant volume) while for drained loading which is accompanied by volumetric
strains a value of v = 02 may be used If we adopt the reference modulus as E then the
interrelationships with the other elastic moduli are given by
a Reference Stiffness 119864prime = drained Youngs modulus A30
119864prime
b Shear Modulus 119866prime = A31 [2(1+120584prime)]
119864prime (1minus120584prime) c Constrained Modulus 119863prime = A32
[(1+120584prime)(1minus2120584prime)]
119864prime
d Bulk Modulus 119870prime = A33 [3middot(1minus2120584prime)]
A-44
Figure A32 Representative stress-strain-strength curve for soils in triaxial compression
For the extreme case when = 0 in fact D = E When = 02 the two moduli are only 10
different D = 11middotE thus the constrained modulus and drained Youngs modulus are often
considered somewhat interchangeable
The resilient modulus (MR) applies to pavement analysis and design most commonly measured by cyclic
triaxial testing under repeated load applications In fact MR is a special case of Youngs modulus that
relates to the small strain stiffness measured in the nondestructive range but has developed permanent
plastic strains after many cycles of loading (Brown 1996 Dehler amp Labuz 2007)
The subgrade reaction modulus is actually a combined soil-structural parameter as its value
depends on the ground stiffness and the size of the loaded element The subgrade modulus is
defined as ks = q where q = applied stress and = measured deflection In terms of elasticity
solutions the deflection of a flexible circle of diameter d is given by = qmiddotdmiddot(1-2)E Therefore
119864prime
119896119904 = A34 [119889middot(1minus1205842)]
which has units of kNm3 or pcf
Table A2 summarizes several selected studies towards the approximate evaluation of these moduli
obtained directly from measured CPT data Note however that the results from in-situ penetrometer
tests generally represent a peak strength as indicated in Figure A32 Thus the measured cone tip
resistance (qt) reflects the top of the stress-strain curve either the undrained strength in clays (su) or the
effective friction angle () of sands or even a strength intermediate between these two values Thus
A-45
Source Soils Studied Reference Modulus
Findings
Kulhawy amp Mayne 1990
8 stiff OC clays Lab constrained modulus
D asymp 825 middot (qt - svo)
Mohammed et al (2000)
Resilient modulus
Abu-Farsakh 2004
7 Louisiana sites
Lab constrained modulus
D asymp 358 middot (qt - svo)
Mayne (2007b)
All soil types Constrained moduli from lab consolidation
First-order estimate D asymp middot(qt - svo)
Robertson (2009)
All soil types Constrained modulus from consolidation
D = m middot (qt - svo)
when Ic gt 22
1 m = Qtn when Qtn le 14
2 m = 14 when Qtn gt 14
when Ic lt 22 then
= 003 middot 10055 Ic + 168 m
Liu et al (2016)
16 clay sites in China
Resilient modulus MR = (146qt 053 + 1355 fs
14 + 236)244
(all in MPa)
Casey et al (2016)
8 clays Triaxial tests for undrained Youngs modulus
Eu(svc)07 = 316 - 23 middot LL
where Eu and svc (MPa)
and LL = liquid limit ()
qt is probably not the best means to obtain the slope of the stress-strain curve Instead the SCPTu that
provides the initial tangent shear modulus (Gmax) from the shear wave velocity (Vs) is a better choice
Table A2 Selected Modulus Relationships with Cone Penetration Test Measurements
A-46
Nonlinear Modulus
The small-strain shear modulus (Gmax or G0) represents the initial stiffness of all soils and rocks It is the
beginning of all stress-strain-strength curves for geomaterials and obtained from elasticity theory
119866119898119886119909 = 1198660 = 120588119905 middot 1198811199042 A35
where Vs = shear wave velocity t = tga = total soil mass density t = total soil unit weight and ga =
gravitational acceleration constant (98 ms2)
From a general viewpoint on stiffness the shear modulus G of soil is defined as the slope of shear stress
versus shear strain G = DDs for a tangent definition and by G = s as a secant definition The shear
modulus is related to its associated Youngs modulus E = 2G(1+v) Both moduli are in fact highly
nonlinear ranging from a maximum value at the small-strain stiffness (Gmax = G0 = t middotVs2 where t =
total soil mass density and Vs = shear wave velocity) to intermediate G values at medium strains (asymp 1)
to low values at peak strength As such a variety of algorithms and formulae have been developed to
represent either a partial range or the full stress-strain-strength behavior of soils over a range of
interests (eg Mayne 2005) In these formulations a variable number of input parameters may be
required in order to produce a stress-strain-strength curve
A modified hyperbolic form suggested by Fahey amp Carter (1993) has favorable attributes in that the
modulus reduction factor can be established with only a single variable This allows the initial stiffness
to the be small-strain stiffness (G0) that is reduced in terms of level of mobilized strength eg max =
qqmax = 1FS which is simply the reciprocal of the calculated factor of safety (FS) The magnitude of
secant shear modulus (G) corresponding to the particular level of loading is given by
119866 = 119872119877119865 middot 1198660 = (1198661198660) middot 1198660 A36
A-47
where the MRF = modulus reduction factor determined as
1 minus (120591 )119892 119872119877119865 = (1198661198660) = frasl A37 120591119898119886119909
and g = empirical exponent term Figure A33 shows the relationships for normalized shear stress ()
versus shear strain (s) and corresponding normalized secant shear modulus (G = s) with shear strain
over a range of exponent values 01 le g le 10 For a discussion of tangent G please see Fahey amp Carter
(1993)
Using laboratory data from resonant column-torsional shear tests and triaxial specimens with local
strain measurements for Gmax reference values modulus reduction curves for a selection of sands and
clays are presented in Figure A34 The data indicate that the exponent g falls within the range 02 lt g lt
05 for many soils tested drained and undrained A mean value of g = 03 is recommended for
preliminary site investigations and designs until additional information can be obtained
The value of max is the shear strength of the soil generally taken as either (1) drained (c = 0) max =
svo middot tan or (2) undrained where max = su = undrained shear strength as discussed earlier Thus each
shear stress () can be associated with its shear modulus (G) and the relevant shear strain is found from
s = G This allows for generation of nonlinear stress-strain-strength curves at all depths from SCPTu
data in clays sands and mixed soil types
A-48
Modified Hyperbola (Fahey amp Carter 1993 Canadian Geotech J) Coefficient Term f = 1
``
0
01
02
03
04
05
06
07
08
09
1
000001 00001 0001 001 01 1
No
rmal
ized
GG
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
0
01
02
03
04
05
06
07
08
09
1
0 01 02 03 04 05 06 07 08 09 1
No
rmali
zed
GG
ma
x
Mobilized Shear Stress max
g =1
07
05
03
02
01
0
01
02
03
04
05
06
07
08
09
1
00001 0001 001 01 1
No
rmali
zed
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
0
01
02
03
04
05
06
07
08
09
1
0 1 2 3 4 5 6 7 8 9 10
No
rmali
zed
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
Reduction Factor
GGmax = 1- (max)g
Figure A33 Normalized modulus (GGmax) and normalized shear stress (max) versus shear strain (s)
for modified hyperbolic relationship of Fahey amp Carter (1993)
A-49
0
01
02
03
04
05
06
07
08
09
1
0 01 02 03 04 05 06 07 08 09 1
Mo
du
lus
Re
du
cti
on
G
Gm
ax
or
EE
ma
x
Mobilized Strength max or qqmax
NC SLB Sand
OC SLB Sand
Hamaoka Sand (e=0625)
Hamaoka Sand
Toyoura Sand e = 067
Toyoura Sand e = 083
Ham River Sand
Ticino Sand
Kentucky Clayey Sand
Kaolin
Fujinomori Clay
Pietrafitta Clay
London Clay
Vallericca Clay
Pisa Clay
Pisa Clay
Pisa Clay
Pisa Clay
Kiyohoro Silty Clay
Thanet Clay
Exp g = 02
Exp g = 03
Exp g = 04
Exp g = 05
MRF = 1 - (qqmax)g
Figure A34 Modulus Reduction Factors (MRF) with mobilized strength of clays and sands
A113 Coefficient of Consolidation
The rate at which foundation and embankment settlements occur as well as the dissipation of excess
porewater pressures is controlled by the coefficient of consolidation (cv) The magnitude of cv is also
required in the design of vertical wick drains that can be installed in soft ground to expedite the time for
consolidation Using the results of CPTu dissipation tests which measure the rate at which the u2 readings
vary with time the in-situ profile of cv can be evaluated Often the piezocone-interpreted values of cv
are validated by comparison with results from laboratory one-dimensional consolidation tests (eg Abu-
Farsakh MY 2004) Alternatively a better method is to cross-check the values with the measured full-
scale performance of instrumented embankments that are constructed over the ground and the recorded
time-rate of consolidation and settlements can provide the best cv for that site (Abu-Farsakh MY et al
2011)
A-50
Monotonic Dissipation Tests
An illustrative dissipation curve from piezocone testing at IDT State Highway 95 at an embankment and
bridge crossing is presented in Figure A35 After the CPTu sounding was halted at a depth of 512 feet
the recorded decay of porewater pressures was observed to be monotonic with time until the dissipations
were ended at 1000 seconds During that time the u2 readings decreased from 115 psi to 47 psi At this
location the depth to the groundwater table is about 16 feet thus the calculated equilibrium water
pressure is 15 psi Commonly a characteristic time for piezo-dissipations is taken at 50 degree of
consolidation although other degrees may be adopted By evaluating the value of u2 at 50 as shown in
Figure A35 the characteristic t50 = 404 s can be obtained
0
20
40
60
80
100
120
1 10 100 1000 10000
Pore
wat
er P
ress
ure
u2
(psi
)
Time (s)
CPTU-02E-1 z = 1560 m
uo (hydrostatic)
u2 (initial) = 115 psi
u2 (final projected) = u0 = 15 psi
u2 at 50 = frac12(115+15) = 65 psi
Idaho DOT State Route 95Dissipation at 512 feet
t50 = 404 seconds
Time to reach50 consolidation
A-51
Figure A35 Piezo-dissipation at Sandpoint Bridge Crossing Idaho for depth z = 512 feet illustrating the
evaluation of characteristic time t50
It is also common to plot normalized excess porewater pressures relative to the measured initial Du2
that is obtained during penetration at the constant rate of 20 mms ie DuDui versus time Figure A36
shows a set of piezo-dissipation records for a bridge and embankment site in southern Louisiana
reported by the LTRC While the porewater pressure axis is arithmetic the time scale is often plotted on
either logarithmic or square root scales In either case the t50 is then found from the measured
dissipation curve when DuDui = 050
A-52
Figure A36 Set of monotonic piezo-dissipations taken at various depths for Courtableau Bridge
Site on Louisana State Highway 103 (Abu-Farsakh et al 2011)
Dilatory Dissipation Curves
In a number of situations a monotonic type dissipation does not occur on the onset but instead a dilatory
type curve is observed Here after stopping the push of the penetrometer the recorded porewater
pressures initially begin to rise and eventually reach a peak value then afterwards follow a phase where
the Du values decrease to hydrostatic (Figure A37) A full solution is available for both cases (Burns amp
Mayne 2002) Dilatory responses are most often associated with overconsolidated clays and silts
although occasionally are observed in fine-grained soils with low OCRs
The selection of the characteristic t50 value may found by using a square root of time plot to find the initial
u2 value by projection of the post-peak dissipatory data back to the ordinate axis as shown by the example
in Figure A38 In this example the initial u2 = 480 kPa and then the same procedures in Figure A35may
apply
A-53
Figure A37 Monotonic and dilatory porewater dissipation responses in soils
(after Burns amp Mayne 1998)
Figure A38 Method for obtaining t50 from dilatory dissipation curves
(Schneider amp Hotstream 2010)
A-54
Interpretation of Dissipation Test Data
There are a number of different solutions available for the interpretation of the coefficient of
consolidation (cv) from the piezocone dissipation curves For monotonic type response the strain path
method (SPM) developed at Oxford University (Teh amp Houlsby 1991) is well-recognized while the Georgia
Tech solution by Burns amp Mayne (2002) is based on spherical cavity expansion and critical state soil
mechanics (SCE-CSSM) and handles both monotonic and dilatory curves For both solutions a simplified
procedure can be recommended that relies on the aforementioned t50 value obtained from the measured
field data as given in Table A3 The units for cv are in cm2s as shown or equivalent
Table A3 Recommended procedures for calculating cv from t50 obtained in dissipation tests
Method Equation for coefficient of
consolidation cv
RemarksNotes Eqn
No
SPM
(Teh amp
Houlsby
1991) 50
2)(2450
t
Iac
Rc
v
t50 = measured time to reach 50
dissipation
IR = Gsu = undrained rigidity index
G = shear modulus
su = undrained shear strength
ac = penetrometer radius (ac = 178
cm for 10-cm2 cone ac = 220 for a
15-cm2 size)
A32
SCE-CSSM
(Burns amp
Mayne 2002) 50
7502 )()(0300
t
Iac Rc
v
A33
A-55
Evaluating rigidity index
Both the SPM and SCE-CSSM solutions require an estimate of the in-situ rigidity index (IR = Gsu) of the
soil If the results of SCPTU are available Krage et al (2014) have derived an expression for IR that
depends upon the small-strain shear modulus net cone tip resistance and effective overburden stress
250750
max50
8111
vonet
Rq
GI
s A38
where consistent units are input for Gmax qnet and svo The value of Gmax is obtained from B29 using the
measured shear wave velocity (Vs) and unit weight evaluated from B7
If the shear wave velocity is not measured it can be estimated from the CPT data A number of
methods have been reviewed for CALTRANS by Wair et al (2012) including one that is applicable
to sands silts and clays of low sensitivity that are inorganic and uncemented
119881 (119898119904) = [101 middot 119897119900119892(119902119905) minus 114]167 middot (100 middot 119891119904119902119905)03 A39 119904
where qt is in units of kPa (Hegazy amp Mayne 1995) An alternative relationship is given by Robertson amp
Cabal (2015)
A114 Hydraulic Conductivity
The hydraulic conductivity is a parameter that expresses the flow characteristics of the soil In
geotechnics it is also called the coefficient of permeability (k) and has units of cms or feetday
Through consolidation theory the hydraulic conductivity relates directly to the coefficient of
consolidation
A-56
A40 D
ck wv
where w = unit weight of water and D = constrained modulus
Therefore one approach to evaluating k can be from the site-specific cv obtained from the
dissipation tests using an estimate of D from one of the relationships given in Table A2
An approach developed for soft normally-consolidated soils that uses t50 directly to assess the magnitude
of k is presented in
Figure A39 (Parez and Fauriel 1988) The mean trendline through the wedges of soil type can be
approximated by
251
50 (sec)251
1)(
tscmkh
A41
A-57
Figure A39 Hydraulic conductivity versus dissipation time for 50 consolidation
(Parez amp Fauriel 1988)
A-58
APPENDIX B
List of Figures
Figure B1 Conventional methods vs direct CPT approach to shallow foundation response B-3
Figure B2 Direct bearing capacity relationship for embedded square and strip footings situated on clean
sands (after Schmertmann 1978) B-6
Figure B3 Direct CPT bearing factor Rk as function of footing width B and embedment depth from finite
element analyses (Tand et al 1995) B-6
Figure B4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio and sand
consistency (Eslaamizaad amp Robertson 1996) B-7
Figure B5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp Salgado (2005) in
terms of footing size relative density and base settlement B-8
Figure B6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and normalized
cone tip resistance (after Eslami amp Gholami 2005) B-8
Figure B7 Direct relationship between ultimate bearing stress in clays and measured cone tip resistance
(Schmertmann 1978) B-10
Figure B8 Direct relationship between ultimate foundation bearing stress in clays and net cone tip
resistance (after Tand et al 1986) B-11
Figure B9 Measured load-displacements for each of the five large footings at TAMU (Briaud amp Gibbens
1994) sand site B-12
Figure B10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU (Briaud amp
Gibbens 1994) footings B-13
Figure B11 Characteristic stress versus square root normalized displacements for TAMU footingsB-15
Figure B12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sandB-16
Figure B13 Summary of sand formation factors with corresponding CPT resistancesB-19
Figure B14 Normalized foundation stresses vs square root of normalized displacement for spread
footings on sand (modified after Mayne et al 2012) B-20
Figure B15 Definitions of capacity from three types of observed foundation load test responses
(modified after Kulhawy 2004) B-21
Figure B16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footingsB-22
B-1
Figure B17 Measured load vs displacement curves for 3 shallow foundations on clay at Baytown Texas
(Stuedlein amp Holtz 2010)B-25
Figure B18 Characteristic stress vs square root of normalized displacement for all three foundations at
Baytown site (Stuedlein amp Holtz 2010) B-26
Figure B19 Normalized foundation stress vs square root of normalized displacements B-27
Figure B20 Applied foundation stress vs CPT-calculated stress for 67 footingsB-28
Figure B21 Applied footing stress vs CPT calculated stress on logarithmic scales B-29
Figure B22 Summary graph and equations of direct CPT method for evaluating the vertical B-30
Figure B23 Foundation soil formation parameter hs versus CPT material index Ic B-32
Figure B24 Bearing capacity ratio qmaxqnet versus CPT material index IcB-33
Figure B25 Influence factor for rectangular foundations from elastic theory solution (Mayne amp
Dasenbrock 2017) B-34
List of Tables
Table B1 Direct CPT methods for bearing capacity of footings on clean sandsB-4
Table B2 Direct CPT methods for bearing capacity of footings on clays B-9
Table B3 Summary of large footings soil conditions and reference sources of database B-17
Table B4 Sources for shallow foundation performace B-34
B-2
B1 DIRECT CPT METHOD FOR SHALLOW FOUNDATIONS
B11 Introduction
The analysis of shallow foundations on soils is classically handled as a two-step and two-part set of
calculations involving (a) bearing capacity commonly reliant on limit plasticity solutions and (b)
settlement or more appropriately termed displacements that are assessed via elastic continuum theory
The utilization of cone penetration tests (CPT) provides the necessary geotechnical data for the site-
specific information on the subsurface conditions at the project site including the geostratigraphy and
evaluation of input parameters This is still a viable approach where the CPT readings are interpreted to
give the soil unit weight (γt) effective friction angle (ϕ) undrained shear strength (su) preconsolidation
stress (σp) and elastic moduli (D and E) for analysis
An alternative approach is the use of CPT results to provide direct assessments of bearing capacity
andor settlements A comparison of these two distinctly different and alternate paths is depicted in
Figure B1
Figure B1 Conventional methods versus direct CPT approach to shallow foundation response
A number of available direct CPT approaches for shallow footings are reviewed within this report
Moreover as the entire load-displacement-capacity response of shallow foundations to loading occurs
as a continuous nonlinear phenomenon a single direct CPT solution is presented for the behavior of
B-3
footings on sands (drained) and clays (undrained) The methods discussed herein are specifically to
address the general case of vertical loading of shallow foundations Additional more complex situations
that consider load eccentricity moments inclined forces sloping ground and other facets may be
handled using well-established procedures that are discussed elsewhere (eg Vesić 1975 Kulhawy et
al 1983)
This report focuses on foundations on granular soils andor soils exhibiting drained behavior since
Federal Highway Administration (FHWA) studies have concluded that less than 1 of shallow
foundations for highway bridges are placed on clay soils (Paikowsky et al 2010) One reason for this is
because soft-firm intact clays require a more complicated process that assesses both a short-term
analysis (undrained) as well as long-term analysis (drained) thus requiring undrained bearing capacity
undrained distortion displacements drained bearing capacity and drained primary consolidation
settlements
B12 Direct CPT Methods for Bearing Capacity of Footings on Sands
Classical bearing capacity solutions are based most often in limit plasticity theory albeit can be
formulated from limit equilibrium cavity expansion and numerical methods Because the limit
plasticity solutions are prevalent and adopt total stress analysis they are usually developed as either
fully drained (Δu = 0) or fully undrained (ΔVV = 0) where Δu equals excess porewater pressure and
ΔVV equals volumetric strain Paikowski et al (2010) provides an extensive review of various solutions
There are a number of direct CPT methods for the evaluation of the bearing capacity of footings and
shallow foundations Table B1 lists a variety of these direct CPT approaches for footings on sands In
the direct approach a one-step process is used to scale the measured cone penetrometer readings (ie
measured cone tip resistance (qc) total cone tip resistance (qt) sleeve friction (fs) andor measured
porewater pressure u2)) to obtain the ultimate bearing stress (qult) of the foundation
Table B1 Direct CPT methods for bearing capacity of footings on clean sands
Method Surface Footing Remarks and Embedment Notes
Meyerhof (1956) qult = qc (B12) middotcw qult = qc (B12)(1+DeB)cw
Note for silty sand reduce by 05
cw = 10 dry or moist sand cw = 05 submerged sand
Meyerhof (1974) qult = qc (B + D)40 with stresses in tsf
Presumably dry sands B = fdn width (feet) and D = depth (feet)
Schmertmann (1978)
NA (applies to embedded footings)
Square qult =055σatm(qcσatm)078
Strip qult =036σatm(qcσatm)078
See Figure A2
Embedment applies De gt 05(1+B) for Blt1m De gt 12 m for B gt1m
B-4
Canadian qult = Rk0middotqc Applied to FS = 3 Geotech Society where Rk0 = 03 where FS = factor of (CFEM 199 2) safety
Tand et al (1995) NA qult = Rkmiddotqc + σvo See Figure A3 where Rk = fctn(De B)
Frank and Magnan (1995)
qult = Rk0middotqc where Rk0 = fctn(sand
consistency) Loose Rk0 = 014
Medium Rk0 = 011 Dense Rk0 = 008
qult = Rk1middotqc + σvo where Rk1 = function (De B L amp
sand consistency) Factor Rk1 = Rk0[1+Rk206+04(BL) (DeB)]
and Rk2 = 035 (loose) 050 (medium) and 085 (dense)
Loose sand qc lt 5 MPa
Medium sand 8 MPa lt qc lt 15MPa
Dense sand qc gt 20 MPa
Eslaamizaad amp Robertson (1996)
qult = KΦmiddotqc See Figure A4
See surface footing equation KΦ = function (BDe shape and density)
Lee amp Salgado (2005)
qbL = βbcmiddotqc(AVG)
where qc averaged over distance B
Not addressed See Figure A5 for factor βbc = fctn(B DR
K0 and sB)
beneath the base
Eslami amp Gholami (2005
2006)
See embedded footing solution
qult = Rk1middotqc where Rk1 = function (ratio DeB
and normalized qcσvo) See Figure A6
Measured qc and qcσvo are geometric means over 2B deep
beneath footing
Robertson amp Cabal (2007)
qult = KΦmiddotqc with KΦ = 016
See surface footing equation
Briaud (2007) qult = KΦmiddotqc with KΦ = 023
Based on full-scale tests at Texas AampM
Mayne amp Illingworth
(2010) qult = 018 qc-Mean
Note qc -Mean is averaged CPT cone resistance over depth of
influence z = 15 B deep
Based on 30 footing load tests on 12
sands
Lehane (2013) qult = 016 qc-AVE Note qc-AVE is averaged CPT cone resistance within z (m) = [B (m)
]07
Based on 47 load tests
Table B1 Continued
Notes B = minimum footing width (or diameter) L = footing length De = embedment depth cw = water
table correction σvo = total overburden stress at bearing elevation and σvo = the effective vertical stress
B-5
Figure B2 Direct bearing capacity relationship for embedded square and strip footings
situated on clean sands (after Schmertmann 1978)
Figure B3 Direct CPT bearing factor Rk as function of footing width B and embedment depth
from finite element analyses (Tand et al 1995)
B-6
Figure B4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio
and sand consistency (Eslaamizaad amp Robertson 1996)
B-7
Figure B5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp
Salgado (2005) in terms of footing size relative density and base settlement
Figure B6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and
normalized cone tip resistance (after Eslami amp Gholami 2005)
B13 Direct CPT Methods for Bearing Capacity of Footings on Clays
For direct CPT evaluations of foundation bearing capacity in clays Table B2 summarizes some of the
well-documented methods that are available Generally these assume that the loading is fast enough
and the permeability of the clay is also sufficiently low such that an undrained (ie constant volume)
condition is maintained This is fine for short term loading of foundations particularly for soft to firm
intact clays However the undrained case is not permanent Given sufficient time excess porewater
pressures in the clay will eventually dissipate and hydrostatic conditions will return to equilibrium
During this dissipation phase primary consolidation will occur that results in drained settlements Also
a drained bearing capacity condition will prevail As the CPT is advanced at a constant rate of 20 mms
the recorded readings normally constitute undrained behavior from a direct measurement viewpoint
Thus direct CPT methods are normally focused at an undrained foundation response Nevertheless the
drained footing case can be addressed using CPT results via use of conventional analysis approach and
effective stress limit plasticity solutions that require piezocone penetrometers with porewater pressure
measurements
B-8
B
D)
L
B4060(3501320kc
Many of the earlier solutions were based on data from mechanical penetrometers andor older electric
cones that measure the uncorrected tip resistance (qc) It is now well-recognized that this measurement
must be updated to the corrected cone resistance (qt) because of porewater pressure effects that act on
the mechanics of the load cell and geometry of the particular penetrometer design The results are
particularly significant in clays silts and mixed soils that are soft to firm to stiff where excess porewater
pressures are recorded
Table B2 Direct CPT methods for bearing capacity of footings on clays
Method Footing Comments NotesRemarks
Meyerhof (1974) qult = αbcmiddotqc
where 025 le αbc le 050
Applicable to saturated insensitive clays under short-term loading
Cone tip resistance
measured by
mechanical CPT
Trofimenkov (1974)
Approximation
qult asymp (qc33)09
(Assuming FS = 3)
Strip footings on clays and sandy clays 06m le B le 15m 1m le zemb le 25m
Mechanical CPT
with qc and qult in
kgcm2
Schmertmann (1978)
qult = function (qc and foundation shape)
Relationship shown in Figure A7 for square and strip footings
Cone tip resistance
measured by
mechanical CPT
Tand et al (1986)
qult = Rkmiddot(qc - σvo) + σvo
Factor Rk obtained from Figure A8 accounts for surface and deep footings Separate Rk factor for intact and jointed clays
= (qc1middotqc2)05 qc
where qc1 is geometric mean from bearing elevation to 05B deeper and qc2 is geometric mean from 05B to 15B beneath foundation base
Mix of data from
mechanical CPT qc
and electric CPT qc
LCPC Method
(Frank amp Magnan 1995)
Footing a surface
qult = kc middotqc + σvo
where kc = 032
Embedment case
Note Methods above generally consider undrained bearing capacity
B-9
Figure B7 Direct relationship between ultimate bearing stress in clays and measured cone tip
resistance (Schmertmann 1978)
B-10
Figure B8 Direct relationship between ultimate foundation bearing stress in clays and net
cone tip resistance (after Tand et al 1986)
B14 Direct CPT Assessments of Settlement and Displacements of Shallow Foundations
Methods for evaluating the magnitude of foundation displacements (s) can be found using
elastic theory subgrade reaction methods spring models and numerical simulations The
classical approach to footing settlement calculations on sands is to utilize elastic theory
solutions (eg Poulos amp Davis 1974) which take on the form
119902∙119861∙119868∙(1minus1205842) 119904 = B1
119864119904
where q = applied footing stress and I = displacement influence factor from elasticity theory
The value of I depends upon foundation geometry footing rigidity embedment depth variation of soil
modulus with depth compressible layer thickness and other factors as discussed by Mayne amp Poulos
(1999) For instance for a flexible circular footing of diameter B resting on an infinitely deep soil
formation having a constant modulus Es with depth the influence factor I is equal to 1 for the center
point The results of in-situ field tests such as pressuremeter tests (PMT) standard penetration tests
(SPT) cone penetration tests (CPT) flat dilatometer tests (DMT) and other methods can be used to
ascertain the input value of soil modulus Es
Alternatively direct CPT methods have been investigated to provide a one-step assessment of
foundation displacements Selected methods for direct CPT evaluation of shallow foundation
settlements on granular soils is include DeBeers amp Martens (1957) Meyerhof (1965) DeBeer (1965)
Thomas (1968) Schmertmann (1970) Meyerhof (1974) Berardi amp Jamiolkowski amp Lancellotta (1991)
Robertson (1991) Lutenegger amp DeGroot (1995) Lehane amp Dougherty amp Schneider (2008) Gavin amp
Adekunte amp OrsquoKelly (2009) Mayne amp Illingworth (2010) and Uzielli amp Mayne (2012)
B141 Large Scale Footing Load Tests
The measured load-displacement response of shallow foundations is conducted in the field using either
stepped loading procedures or continuous rate of displacement methods Results from the well-
documented test program at the Texas AampM University (TAMU) site (Briaud amp Gibbens 1999 Briaud
2007) for five spread footings on sand are shown in Figure B9 Each footing clearly shows a nonlinear
B-11
behavior to loading over the range of testing This site is one of the national geotechnical
experimentation sites (NGES) funded by the National Science Foundation (NSF) FHWA and American
Society of Civil Engineering (ASCE)
Figure B9 Measured load-displacements for each of the five large footings at TAMU (Briaud
amp Gibbens 1994) sand site
B142 Generalized Direct CPT Method for Footing Response on Soils
The use of applied foundation stress versus normalized displacement curves (q vs sB) is generalized to
footings on sands silts intact clays and fissured clays A square root plotting of the normalized
displacements (sB) permits a single parameter characterization for each specific soil type that in turn
correlates with the net cone tip resistance The generalization is based on a statistical review of data
from large foundations (B gt 05m) involving 70 full-scale load tests with available CPT data For fine-
grained soils the data are from electronic piezocone tests where the proper correction from qc to qt has
been obtained to allow the best possible results The methodology permits an evaluation of the load-
displacement-capacity response of shallow square footings based on CPTs performed in the four types
B-12
of ground conditions For the TAMU footings the characteristic stress-normalized displacement
response is shown in Figure B10 where all five foundations can be represented by a single curve
Figure B10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU
(Briaud amp Gibbens 1994) footings
B143 Characteristic Stress-Displacement Curves
Fellenius amp Altaee (1994) recommended the concept of a characteristic stress (q) vs normalized
displacement (sB) curve for foundation response on a given soil formation later supported by Decourt
(1999) Briaud amp Gibbens (1999) Lutenegger amp Adams (2003) and Briaud (2007) In this approach the
measured load (Q) vs displacement from individual footings of various sizes that rest on the same soil
conditions collapse to a unified relationship given in Equation B2
119939119943 119956 119954119938119953119953119949119946119942119941 = 119938119943 (119913
) B2
B-13
where af and bf are empirical fitting coefficients (Decourt 1999 Uzielli amp Mayne 2011 2012)
In a review of measured load tests data from large spread footings situated on different sands it has
been suggested that Equation B2 can be reduced to a single parameter expression by use of square root
plotting where the exponent bf is equal to 05 (Mayne amp Illingworth 2010 Mayne et al 2012)
119956 119954119938119953119953119949119946119942119941 = 119955119956radic
119913 B3
where rs = fitted soil parameter from best fit line regression In the case of sands the coefficient rs
represents the supporting soil conditions including particle size relative density and fines content as
well as geologic origin aging and overconsolidation effects
The results from the five TAMU footings are presented in this format in Figure B11 that shows the
formation factor rs = 486 MPa for this sand deposit This single coefficient can be used to express the
nonlinear load versus settlement of any size footing on this sand site Moreover one criterion from
Eurocode for bearing capacity that is used by the European community identifies that stress (or load)
which corresponds to (sB) = 10 That is when the displacement equals ten percent of the footing
width then the capacity has been reached Therefore adopting this criterion for footings on sands
the ultimate bearing stress (qult) for footings on sand can be taken when (sB) equals 010 or when
square root of (sB) equals 0316 In that case this gives qult equal to 0316 middot rs For the TAMU site this
gives qult equal to 0316 times (486) which equals 153 MPa as illustrated by Figure B12
B-14
Figure B11 Characteristic stress versus square root normalized displacements for TAMU
footings
B-15
Figure B12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sand
B15 Footing Database
A database of spread footings and large plates was assembled which considers only full-size shallow
foundations (05 le B le 6 m) that rest on sands silts and clays Sites for the foundations were also subjected to CPT The inclusion solely of large footings are significant because results from small-scale
model footings exhibit scale effects In fact experimental centrifuge work and numerical simulation
studies have shown that bearing capacity factors are size-dependent especially for factor Nγ decreasing
with footing width B (Kimura et al 1985 Cerato amp Lutenegger 2007 Mase amp Hashiguchi 2009) A
direct approach based on full-size footings provides a more reliable means for foundation evaluation
The compiled database is given in Table B3 with a total of 70 large footings and plates These include a
listing of 34 foundations on 13 sands 11 footings on 4 silt deposits 13 footings or large plate load tests
on 6 intact clays and 12 foundations on 5 fissured clays Most of the footings were square or nearly
square (80) while the remaining were circular The largest foundation was a mat 5 m by 14 m in plan
loaded by stacking Kentledge concrete blocks to cause a bearing failure in the underlying soft clay For
sands the largest footing consisted of the 6-m square concrete pad In terms of equivalent square
footings mean footing sizes were 155 m (sands) 114 m (silts) and 126 m (clays)
B-16
Additional details on the individual load tests and site conditions are given elsewhere (Mayne 2009
Mayne amp Illingworth 2010 Uzielli amp Mayne 2011 2012 Mayne et al 2012 Mayne amp Woeller 2014)
Table B3 Summary of large footings soil conditions and reference sources of database
Sand Site Location Soil Conditions Footings Numbers
Shapes and Sizes
ReferencesSource
College Station
Texas Pleistocene sand 5 Square 1 15 25 33 m Briaud amp Gibbens (1999)
Kolbyttemon Sweden Glaciofluvial sand 4 Rect B = 06 12 17 24 m
Bergdahl et al (1985 1986)
Fittija Sweden Glaciofluvial sand 3 Rect B = 06 17 24 m Bergdahl et al (1984 1985)
Alvin West Texas Alluvial sand 2 Circular D = 235 m Tand et al (1994)
Alvin East Texas Alluvial sand 2 Circular D = 22 m Tand et al (1994)
Perth Australia Silceous dune sand 4 Square B = 05 and 10 m
Lehane (2008)
Grabo T1C Sweden Compacted sand fill
1 Square B = 046 m Long (1993)
Grabo T2C Sweden Compacted sand fill
1 Square B = 063 m Long (1993)
Grabo T3C Sweden Compacted sand fill
1 Square B = 080 m Long (1993)
Labenne France Dune sand 6 Square B = 07 and 10 m
Amar et al (1998)
Green Cove Florida Brown silty sand 1 Circular D = 182 m Anderson et al (2006)
Durbin South Africa
White fine sand 1 Square B = 609 m Kantley (1965)
Porto Portugal Residual clayey sands
1 Circular D = 12 m and 1 plate
Viana da Fonseca (2003)
Jossigny France Soft clayey silt 2 Square B = 1 m Amar et al (1998)
Tornhill Sweden Glacial Baltic till 3 Square B = 05 1 2 m Larsson (2001)
Vagverkel Sweden Stiff medium silt 3 Square B = 05 1 2 m Larsson (1997)
Vattahammar Sweden Brn-Gry layered silt 3 Square B = 05 1 2 m Larsson (1997)
B-17
Table B3 Continued
Clay Site Location Soil Conditions Footings Numbers Shapes and Sizes
ReferencesSource
Baytown Texas Fissured Beaumont 2 plates with d = 076 m Stuedlein amp Holtz (2008)
Belfast Ireland Soft clay silt ldquosleechrdquo
1 Square Pad B = 20 m Lehane (Geot Engr 2003)
Bothkennar Scotland Soft silty clay 2 Square Pads B = 22 24 m
Jardine et al (1995)
Bangkok Thailand Soft to stiff clay 4 Square 067 075 090 10 m
Brand et al (1972)
Haga Norway Stiff OC clay 2 Square Footings B = 10 m
Andersen amp Stenhammar (1982)
Rio Grande Brazil Sandy residual clay 3 Square 04 07 and 10 m
Consoli et al (1998)
Shellhaven England Soft estuarine clay Rectangular 5 m by 14 m Schnaid et al (1993)
Texas City A Texas Coast Fissured Beaumont 3 circular plates d = 058 m
Tand et al (1986)
Texas City B1 Texas Coast Fissured Beaumont 2 circular plates d = 058 m
Tand et al (1986)
Texas City B2 Texas Coast Fissured Beaumont 1 circular plate d = 058 m
Tand et al (1986)
Alvin Texas Texas Coast Fissured Beaumont 3 circular plates d = 058 m
Tand et al (1986)
For the series of footings on sands Figure B13 shows the summary of measure formation factors (rs) for
the 34 foundation load tests Also indicated on the graph are the corresponding mean qc values of the
sands averaged over 15middotB beneath the bearing elevations
Note also in sands that the qt and net resistance (qnet) are all very close to the measured qc since (a)
porewater pressure corrections for the a-net value are negligible in clean sands and (b) overburden
stresses are typically only 1 to 3 of the magnitude of qc This is especially true of shallow depths
Therefore for clean sands qnet is approximately qt which is approximately qc
B-18
Figure B13 Summary of sand formation factors with corresponding CPT resistances
As suggested by Briaud (2007) various in-situ measurements can be used to normalize the results from
cone penetration tests in this case Herein the qc from the CPT soundings in the sands are used to
normalize the footing stress axis as presented in Figure B14 It is evident that the results of the load-
displacement response of all 34 footings on 13 different sands can be captured by the characteristic
stress versus normalized displacement with the inclusion of the site-specific cone tip resistance In this
case the mean trend for all footings and sand sites indicates a simple expression shown in Equation B4
119954119943119952119952119957119946119951119944 119956 Clean quartz-silica sands = 120782 120787120790120787radic B4
119954119940 119913
B-19
Figure B14 Normalized foundation stresses vs square root of normalized displacement for
spread footings on sand (modified after Mayne et al 2012)
The results of measured force-displacement curves (Q vs s) from footing load tests are nonlinear and
thus the definition of capacity must be addressed Kulhawy (2004) identified three main curve types
that occur during load testing depicted in Figure B15 The most common response (Type A) shows no
clear peak and is observed in sands and silts as well as slow loading of insensitive clays and fissured
clays In this case ldquocapacityrdquo can be defined by the Eurocode corresponding to the load (or stress) when
sB=10 (Amar et al 1998) For foundations situated on many clays a plateau capacity is reached as
depicted as Type B response In rare cases where sensitive clays or structured soils are encountered a
Type C relationship occurs whereby the load-displacement curve reaches a peak followed by strain
softening In the one case observed in this study for Type C loading (Haga clay) the corresponding
ldquopeak capacityrdquo was reached at a pseudo-strain of sB = 4 as shown in Figure B16 followed by some
strain softening In the cases of soft clays at Belfast and Bothkennar a plunging type (plateau failure)
was achieved at around sB=7
B-20
Figure B15 Definitions of capacity from three types of observed foundation load test
responses (modified after Kulhawy 2004)
B-21
Figure B16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footings
In the case of fine-grained soils that develop excess porewater pressure during cone penetration the
use of the net total cone resistance (qtnet) which equals the cone tip resistance minus the total stress
must be considered (Lunne et al 1997) As noted earlier the correction of CPT data in sands is minimal
because of the low value of penetration porewater pressures and the fact that total overburden stress is
small relative to the cone resistance especially for shallow soundings Thus the use of qtnet may be
substituted for qc into the trend of Figure B1
B16 Undrained and Drained Capacity
The footings on sands typically show drained response with no excess porewater pressures developed
during loading For the foundations situated on silts analyses by Larsson (1997 2001) concluded that
these too show drained conditions For shallow foundations on clays the entire range of drainage cases
are possible ranging from undrained to partially-drained to fully-drained depending upon the applied
rate of loading relative to the permeability of the geomaterial If sufficient data were available for each
of the clay sites then the recent evaluation of drainage condition could be assessed using normalized
velocity criteria (Chung et al 2006) However this was not possible with the current data set Instead
B-22
the test loadings of footings on clays have essentially been assumed to primarily occur under undrained
conditions following recommendations by Tand et al (1986) For this case a limiting bearing stress can
be calculated from bearing capacity analyses and related directly to the CPT readings
From classical bearing capacity calculations using limit plasticity solutions the ultimate foundation stress
is shown in Equation B5
= 119925119940 ∙ 119956119958 B5 119954119958119949119957
where Nc equals the bearing factor for constant volume (514 for strip and 614 for square or circular
plan) and su equals the undrained shear strength of the clay For the case of soft to firm clays of low to
medium sensitivity the operational strength can be related to the preconsolidation stress (Mesri 1975
Jamiolkowski et al 1985) as shown in Equation B6
prime 119956119958 = 120782 120784120784120648119953 B6
Finally the effective preconsolidation stress has been linked directly to the net cone resistance (Chen amp
Mayne 1996 Demers amp Leroueil 2002) as shown in Equation B7
prime 120648119953 = 120782 120785120785(119954119957 minus 120648119959119952) B7
Combining Equations (B5) (B6) and (B7) results in direct expressions for undrained bearing capacity on
intact clays from CPT results as shown in Equations B8 and B8-2
Strip footing 119954119958119949119957 = 120782 120785120789120785(119954119957 minus 120648119959119952) B8
B-23
Squarecircle 119902119906119897119905 = 0445(119902119905 minus 120590119907119900) B8-2
Equation (B8-2) is in excellent agreement with Tand et al (1986) database study for the case of intact
clays and footings with no embedment Their study also provided a lower relationship for foundations
situated on jointed clays specifically recommending that the bearing stress not exceed 030 qtnet
Review of the available data in Figure 3 shows this to be rather conservative and a more realistic value
may be taken at 040 qtnet for fissured clays
B161 Normalized Undrained Footing Response
The characteristic stress vs normalized displacement curves for footings on clays exhibiting undrained
conditions can be verified using a relatively recent case study on Beaumont clay by Stuedlein amp Holtz
(2010) The Baytown Texas site was investigated using an exploration program of soil borings
piezocone soundings and laboratory testing The field testing included a large square footing (B = 276
m) and two small circular plates (D = 076 m) The measured load-displacement responses for all three
foundations are presented in Figure B17 Of course the large footing carried considerably higher loads
than the two small plates on the same soil deposit Yet using the aforementioned characteristic stress
(q) vs square root of normalized displacement (sB) essentially gave a unique relationship for all 3
foundations as presented in Figure B18 In this case utilizing the form of Equation B2 the
characteristic coefficient rs is 177 MPa for this deltaic clay formation
B-24
Figure B17 Measured load vs displacement curves for 3 shallow foundations on clay at
Baytown Texas (Stuedlein amp Holtz 2010)
B-25
Figure B18 Characteristic stress vs square root of normalized displacement for all three
foundations at Baytown site (Stuedlein amp Holtz 2010)
B17 General Direct CPT Method
In a manner analogous to the normalization of applied foundation stresses shown in Figure B14 for
footings on sands a generalized direct CPT method for shallow foundations on different soils can be
made in Equation B9
120782120787 119956 119954 = 119945119956 ∙ 119954119957119951119942119957 ∙ (119913
) lt 119954119940119938119953119938119940119946119957119962 B9
where qcapacity is equal to the foundation bearing capacity of the ground and hs is equal to the empirical
fitting term that depends on soil type Specifically hs equals 058 for sands 112 for silts 147 for
fissured fine-grained soils and 270 for clays A summary of the normalized stress versus the normalized
displacement curves for these four soil types is presented in Figure B19 where the data fitting to obtain
the hs parameter on clays are limited to (sB) lt 004 corresponding to undrained loading The data on
B-26
silts and sands are considered fully drained whereas the fissured clay subset may be partially drained to
undrained The statistical measures for obtaining the fitted parameter hs are quite good as shown in
Figure B20 with the coefficient of determinations (r2) values of 0947 for sands 088 for silts 0935 for
fissured and 0925 for intact clays Additional statistical evaluations on the database are provided by
Uzielli amp Mayne (2011)
Figure B19 Normalized foundation stress vs square root of normalized displacements
B-27
Figure B20 Applied foundation stress vs CPT-calculated stress for 67 footings
The same data are shown on Figure B21 in a log-log plot to emphasize the early parts of the load-
displacement responses of these footings The best fit line from regression analyses shows excellent
statistical correlations As detailed earlier the undrained bearing capacity of square or circular footings
on clays can be taken as approximately 040 times qtnet For silts and sands that experience drained
loading with no excess porewater pressures being developed and no clear peak value for capacity the
(sB) =10 criterion can be used In this case Equation 7 can be used directly to obtain stress q equal to
qcapacity when (sB)05 is 0316
An alternate approach is presented in Figure B22 summarizing the direct CPT method for estimating the
load-displacement-capacity of shallow square footings on four soil types The maximum values of soil
bearing capacity (qmax) can be capped at stresses corresponding to a percentage of the average
measured qtnet determined over the depth interval from the foundation bearing elevation to 15middotB
below the foundation In such an approach the foundation bearing capacity can be taken as
(qcapacityqtnet) of 02 for sands 035 for silts 040 for fissured and 045 for intact clays Using the assigned
values of the hs parameter the corresponding cut-off for the general stress-normalized displacement
relationship given by Equation 7 can be established specifically giving corresponding values of the
B-28
normalized displacements which can also be considered as pseudo-strains equal to (sB) max of 4 for
clays 7 for fissured clays 10 for silts and 12 for sands
Figure B21 Applied footing stress vs CPT calculated stress on logarithmic scales
B171 CPT Soil Material Index Ic
The soil behavioral type can be assessed indirectly by CPT using a material index (Ic) as defined by
Robertson (2009a b) shown in Equation B10
119920119940 = radic[120785 120786120789 minus 119949119952119944119928119957119951]120784 + [120783 120784120784 + 119949119952119944119917119955]120784 B10
B-29
Figure B22 Summary graph and equations of direct CPT method for evaluating the vertical
where Qtn equals the stress-normalized cone tip resistance and Fr equals the normalized sleeve friction
calculated from the cone penetrometer readings as shown in Equations B11 and B12 respectably
(119954119957minus120648119959119952)120648119938119957119950 = 119951 B11 119928119957119951 prime (120648119959119952120648119938119957119950)
119943119956 119917119955() = 120783120782120782 ∙ B12 (119954119957minus120648119959119952)
where σatm equals a reference stress equal to atmospheric pressure (σatm = 1 atm asymp 1 bar asymp 100 kPa) In
the initial evaluation the exponent n is set to n = 1 to find the soil behavioral type (SBT) based on a 9-
zonal chart as shown in Figure 23 The value of exponent n varies with soil type ranging from n = 1 in
intact clays and decreasing with increasing grain size to around approximately 075 in silts and
B-30
approximately 05 plusmn 02 in clean sands The appropriate value of exponent n is found by iteration until
conversion using the relationship (Robertson 2009b) shown in Equation B13
prime 120648119959119952 119951 = 120782 120785120790120783 ∙ 119920119940 + 120782 120782120787 ( ) minus 120782 120783120787 le 120783 120782 B13 120648119938119957119950
As the distinction that footings on intact clays subjected to fast loading are behaving primarily under
undrained conditions (ie constant volume) Robertson (2009a) suggested that this occurs when Ic gt 27
while in contrast Ic lt 25 corresponds more or less to drained loading (ie no excess porewater
pressures)
The CPT material index can be used to identify soil type and the corresponding characteristic hs value for
estimating footing load-displacement response In a number of the case studies investigated the results
of the sleeve friction readings were also available to allow the calculation of Ic with depth at these sites
Figure B23 shows the tentative relationship between the values of the hs parameter plotted versus the
corresponding CPT material index with an approximate trend given by Equation B14
120784120785 119945119956 = 120784 120790 minus 120783120787 B14
119920119940 120783+( ) 120784120786
B-31
Figure B23 Foundation soil formation parameter hs versus CPT material index Ic
Soil behavioral type from the Ic ranges given by Robertson (2009b) and are shown in Figure B23 with an
overall general agreement with the known soil classifications In summary the CPT can provide an
evaluation of the load-displacement-capacity response directly via Equation B15 which may be of
interest in mixed soil types such as silty sands sandy silts and the like
120784120785 Footing stress 119954 = 119954119951119942119957 ∙ radic(119956frasl119913) ∙ [120784 120790 minus
)120783120787] B15
120783+(119920119940frasl120784120786
As discussed earlier foundation bearing capacity may be taken by a limiting value of pseudo-strain (sB)
max or by qmax as specified as a percentage of qtnet This definition of bearing capacity can be seen in
comparison to soil type as seen in Figure B24
B-32
Figure B24 Bearing capacity ratio qmaxqnet versus CPT material index Ic
B172 Rectangular Foundations
This same elastic solution from Giroud (1968) for a CPT method for square and circular foundations were
utilized for rectangular foundations The study covered rectangular distortions (AB) ranging from 1
(square) to very long foundations with AB equal to 20 where A represents the foundation length and B
represents the foundation width (Mayne amp Dasenbrock 2017) The influence factor for rectangular
shaped foundations is given in Figure B25 and the expression in Equation B16
= (119912119913)120782120785120786120787 119920119912119913 B16
B-33
Figure B25 Influence factor for rectangular foundations from elastic theory solution (Mayne
amp Dasenbrock 2017)
Including 32 full scale load tests on sands results from an additional 98 shallow foundations have been
reported in previous studies The foundations were primarily square or circular and among these include
large spread footings with measured settlements and bearing capacity Table B4 shows a summary of
the number of footings and footing sizes used in the study The range of length to width ratios varied
from 1 lt (AB) lt 23 with an average value of (AB) equal to 238 The embedment depth to footing
width ranged from 0 lt DeB lt 222 and averaged 046
Table B4 Sources for shallow foundation performace
Source of Data No of
Footings
Mean B
(m)
Max B
(m)
Min B
(m)
Mean A
(m)
Max A
(m)
Min A
(m)
Mayne et al (2012) 32 149 609 046 149 609 046
Lehane (2011) 8 041 027 060 041 060 027
Gifford et al (1987) 17 846 1588 527 1591 3596 701
Jeyapalan amp Boehm
(1986)
13 1022 2892 400 1671 3049 640
B-34
Schmertmann
(1970)
31 945 5608 061 1288 8668 061
Papadopoulos
(1992)
29 870 3600 100 1300 7290 100
Total 130 671 5608 027 1012 8668 027
The solution for shallow foundation settlements given back in Equation B1 combined with
the expression in Equation B16 takes on the form
119954∙119913∙119920119912119913∙(120783minus120642120784) 119956 = B17
119916119956
Combining the direct CPT equation developed from the 32 full scale load tests (Equation B15)
together with the influence factor for rectangular shaped foundations becomes
minus120782120785120786120787 120784120785 119912 Footing stress 119954 = 119954119951119942119957 ∙ radic(119956frasl119913) ∙ [120784 120790 minus
)120783120787] ∙ [ ] B18
120783+(119920119940frasl120784120786 119913
B18 Conclusions
A direct CPT method for square rectangular and circular shallow footings is developed using a database
of 166 full-scale field load tests where the minimum footing size of an equivalent square width of 05 m
is established to avoid scaling issues The use of load vs displacement is generalized by characteristic
stress vs normalized displacement (sB) and further simplified by a square root plotting technique that
captures the ground response in a single parameter that is related to the qtnet Data are grouped
according to four main soil categories sands silts fissured clays and intact clays It is generally believed
that the footings on sands and silts exhibit fully drained behavior while in intact clays an undrained
response occurs under conditions of constant volume
B-35
The summary equation for evaluating the vertical stress-displacement-capacity of square rectangular
and circular footings is given by
minus120782120785120786120787 119956 119912 119954119943119952119952119957119946119951119944 = 119945119956 ∙ 119954119957119951119942119957 ∙ radic ∙ ( ) lt 119954119950119938119961 B19
119913 119913
where the parameter hs also relates directly to Ic The foundation capacity depends upon the mode of
failure as well as drainage conditions and can be taken as a fraction of qtnet where qmaxqtnet is 020 in
sands 035 in silts 040 in fissured clays and 045 in intact clays as shown in Figure 24 Alternatively the
capacity can be defined using a limiting pseudo-strain given by (sB) max of 4 in clays 7 in fissured
10 in silts and 12 in sands
B-36
APPENDIX C
List of Figures
Figure C1 Components of Axial Pile Capacity C-3
Figure C2 Selected alpha relationships as function of the undrained shear strengthC-5
Figure C3 Pile side friction coefficient β in terms of effective friction angle and overconsolidation ratio
C-11
Figure C4 Concept of Direct versus Rational Method for CPT evaluation of axial pile capacityC-14
Figure C5 Soil behavior type using CPT via original UniConeC-19
Figure C6 Soil behavior type using CPT via Modified UniCone C-19
Figure C7 Nine-zone soil behavioral soil type using normalized piezocone parameters C-21
Figure C8 Summary of Modified UniCone Method for direct CPT assessment of axial pile capacity C-22
Figure C9 Elastic continuum solution for axial pile response under compression and tension loadingC-24
List of Tables
Table C1 Alpha Methods for Axial Pile Side Friction where fp = αmiddotsu C-5
Table C2 Beta Methods for Axial Pile Side Friction where fp = βmiddotσvoC-9
Table C3 Selection of Direct CPT Methods for Axial Pile CapacityC-15
C-1
C1 EVALUATING DEEP FOUNDATION RESPONSE FROM CONE
PENETRATION TESTS
C11 Introduction
The axial response of deep foundations includes the load-displacement-capacity and axial load transfer
when driven pilings and drilled shafts are loaded in compression and uplift The use of cone penetration
testing (CPT) especially seismic piezocone tests (SCPTu) are advantageous since they provide
information on the subsurface soils and their geomaterial properties The SCPTu provides at least four
measurements on soil behavior with depth including (a) cone tip resistance qt (b) sleeve friction fs (c)
penetration porewater pressure u2 and (d) shear wave velocity Vs This offers the opportunity to
determine the geostratigraphy unit weight effective overburden stress shear strength stress state
and stiffness of the ground from a single exploratory sounding thus values in economic expediency
and reliability The axial pile capacity can be calculated based on static equilibrium of forces acting along
the sides and base of the pile foundation The displacement of the pile can be ascertained using
elasticity theory from closed-form solutions boundary elements andor finite element analyses Load
transfer occurs along the length of the pile and only a portion of axial forces are transmitted to the base
or toe or tip of the pile These too can be assessed using elasticity solutions
An alternate approach to pile capacity involves the utilization of direct CPT methods available
since the 1970s but in the past 10+ years a number of new and statistically reliable algorithms
have been developed which can be implemented for highway design and construction
C111 Axial Pile Capacity
The axial compression capacity of a single pile foundation is composed of a shaft or side component and
end-bearing component at the base as depicted in Figure 1 For a circular pile the side capacity (Qs) is
determined from the unit side friction (fp) acting along the surface area of the shaft which is As = πmiddotdmiddotL
where d = pile diameter and L = length embedded below grade
If the magnitude of side friction is uniform and constant with depth the side capacity is simply
C-2
119928119956 = 119943119953 middot 119912119956 C1
Moreover however many piles extend through multiple layers and a summation of unit side
frictions acting on various pile segments must be tabulated over the length of the pile as
suggested by Figure C1
Figure C1 Components of Axial Pile Capacity
The unit-end bearing resistance (qb) acts over the base of the pile tip where the area of a circular pile is
given by Ab = πmiddotd24 For piles in compression loading the base capacity is determined from Equation
C2
119876119887 = 119902119887 middot 119860119887 C2
and for piles in tension (or uplift) it is normally taken that Qb = 0
C-3
C112 Pile Unit Side Friction
Several different approaches can be adopted for evaluating the unit pile side friction (fp) prior to full-
scale load testing and construction (Poulos amp Davis 1980 ONeill 2001) The most common types
including the alpha and beta methods as summarized in Tables 1 and 2 respectively In the alpha
method an empirical coefficient (α) is applied to the undrained shear strength (su) of clay soils to
obtain
119891119901 = 120572 middot 119904119906 C3
An illustrative example of alpha curves is shown in Figure C2 A difficulty with the alpha
method is the evolution of many variants and changes to the expressions for its estimation It is
also restricted in its use for specific types of piles in clays and fine-grained soils
C-4
Figure C2 Selected alpha relationships as function of the undrained shear strength
Table C1 Alpha Methods for Axial Pile Side Friction where fp = αmiddotsu
C-5
Reference
Source
Alpha Equation Remarks
Tomlinson
(1957)
2716801102
uu ssAll pile types (steel
concrete timber)
Note su in ksf
Tomlinson
(1957)
2616201102
uu ssConcrete piles
Note su in ksf
McClelland
(1974)
Kerisel α = 07 - 031middotln(su)
Peck α = 10 - 02middotsu
Woodward α = 091middot(su)091
Driven piles in clay
Note su in ksf
Semple
(1980) )1(511
)1(1
OCR
OCR Summary from 9 series of
pile load tests
American
Petroleum
Institute (API
1981)
For suσvo le 035 α = 10
For suσvo gt 080 α = 05
Otherwise α = 1 + 1111middot (035 - suσvo)
Driven steel pipe piles
Tomlinson
(1986)
1 α = 55 (su)-091
2 α = 785 (su)-102
3 α = 605 (su)-100
where 75 lt su (kPa) lt 210
1 Driven piles
2 Bored piles
3 Driven piles in till with L
lt 10middotd
API (1987) 1 α = 10 for su lt 25 kPa Driven piles in clay other
than Gulf of Mexico
C-6
2 α = 125 - 001middotsu for 25 lt su lt 75 kPa
3 α = 050 for su gt 75 kPa
Kulhawy amp
Jackson
(1989)
α = 021+026middot(σatm su) le 1 Analyses of 106 drilled
shaft foundations
Table C1 Continued
API (1989) 05 For suσ vo le 1 α =
su radic frasl prime σvo
minus025 su For suσ vo gt 1 α = 05 ∙ ( frasl prime ) σvo
Driven steel pipe piles
Kulhawy amp
Jackson
(1989)
OCR
OCR
sin50
tan)sin1( sin
Λ = (1-CsCc) asymp 080 for
insensitive clays
Nowacki et al
(1996)
minus05 su For suσvo le 07 α = 05 ∙ ( frasl prime ) σvo
minus02 su For suσvo gt 07 α = 055 ∙ ( frasl prime ) σvo
Driven pile foundations
Kolk and van
der Velde
(1996)
α =09middot FL middot (suσvo)-03 lt 10
FL = [ (L - z)d]-02 = length term
L = pile length
Note su obtained from
UU lab tests
Applicable to driven piles
in clays
C-7
Knappett amp α = 1 for su le 30 kPa Non-displacement piles in
Craig (2012) fine-grained soils α = 116 - (su185) for 30 kPa le su le 150 kPa
α = 035 for su gt 150 kPa
Knappett amp
Craig (2012) α = 055 ∙ 02 40
( ) (LfraslD)
∙ su ( frasl prime σvo
minus03
) Displacement piles in fine-
grained soils
z = depth at point considered
d = outside diameter of pile
Karlsrud et al
(2005 NGI
Method)
α = function (suσvo and plasticity index) Driven pilings
Fleming et al
(2009) 05 minus05 su su For su σvo le 1 120572 = ( frasl prime ) ∙ ( frasl prime )
σvo NC σvo
05 minus025 su su For su σvo gt 1 α = ( frasl prime ) ∙ ( frasl prime ) σvo NC σvo
For driven piles in clay
where (suσvo)NC is the
normalized shear strength
for normally-consolidated
clay
Brown et al
(2010)
α = 0 for 0 lt z le 5 feet
α = 055 for z gt 5 feet and (suσatm) le 15
α = 055-01middot(suσatm -15) for 15le (suσatm) le 25
Drilled Shafts and Bored
Piles
Table C1 Continued
C-8
OCRK )sin1(0
Karlsrud
(2012)
α = function (su σvo and plasticity index) Driven pilings
Note as reported by Karlsrud (2012)
In the beta method the coefficient β is applied to the effective overburden stress (σvo) at the point of
concern along the pile length
prime = 120573 C4 119891119901 middot 120590119907119900
Table C2 Beta Methods for Axial Pile Side Friction where fp = βmiddotσvo
Reference Source Beta Equation Remarks
Burland (1973) β = (1-sinϕ) middot tan ϕ Piles in NC clays
Meyerhof (1976) β = Kmiddottanδ where K = K0 = for NC clays
K = 15middotK0 for OC clays
Poulos amp Davis
(1980)
β = (1-sinϕ) middot tan ϕ middot OCR05 Piles in OC stiff clays
Table C2 Continued
C-9
Kulhawy amp Jackson
(1989)
β = (1-sinϕ) middot tan ϕ middot OCRsinϕ Piles in quartz sands and insensitive
clays
ONeill (2001) β = (1-sinϕ) middot tan θ middot OCRsinϕ Drilled shafts where θ = (δϕ)middotϕ and
δ = interface friction between pile and soil
Fleming et al
(2009)
β = Kmiddottanδ K = lateral stress coefficient and δ =
friction angle between soil and pile
material
Karlsrud (2012) β = fctn (OCR and PI) PI = plasticity index of the clay ()
Mayne and Niazi
(2017)
β = CM middot CK middot K0 middot tanϕ
where K0 = (1-sinϕ) middot OCRsinϕ for soils
that are virgin loaded then unloaded
CM = pile material factor = 1 (drilled
augered) 09 (prestressed or precast
concrete) 08 (timber) and 07
(steel) and CK = pile installation factor
= 09 (bored or augered) 10 (low
displacement eg H-pile or open-end
pipe) and 11 (driven high
displacement eg prestressed
concrete closed-end pipe)
As the beta approach applies to all types of soils (gravels sands silts and clays) and more or
less has remained unchanged since its advent circa 1970 it has been selected for further
discussion herein In the direct form for calculation of unit pile side friction the expression is
given by
119874119862119877119904119894119899120601 prime prime 119891119901 = 119862119872 ∙ 119862119870 ∙ (1 minus 119904119894119899120601prime) middot ∙ 119905119886119899120601prime ∙ 120590119907119900 C5
C-10
where CM is a soil-pile interface friction coefficient and CK = pile installation factor Values of CM are in
the range 07 lt CM lt 10 depending upon pile material (steel wood concrete) and ranges for CK vary
09 lt CK lt 11 depending upon method of installation (auger drill driven) as detailed in Table 2
The value of K0 is limited to the passive stress coefficient which for the simple Rankine case is given by
KP = (1+sin ϕ) (1 - sin ϕ ) Thus there is a maximum value of overconsolidation ratio for which
Equation C5 applies given by OCRlimit = [(1+sin ϕ ) (1 - sin ϕ )2] (1sinϕ)
The corresponding graph in Figure C3 shows the relationship for β in terms of ϕ and OCR
Figure C3 Pile side friction coefficient β in terms of effective friction angle and overconsolidation ratio
C113 Pile Unit End Bearing
C1131 Theoretical Considerations
The evaluation of the end bearing resistance of pile foundations is commonly determined using
limit plasticity theory where
prime Drained loading = C6 119902119887 119873119902 middot 120590119907119900
C-11
Undrained loading 119902119887 = 119873119888 middot 119904119906 C7
where the value of σvo is calculated at the depth z = L and the value of su is the average undrained shear
strength from z = L to a depth z = L + d beneath the pile tip For a circular pile the limit plasticity solution
for undrained loading gives (Vesic 1977) a value Nc = 933 while a deep strip foundation would employ a
value of Nc = 824 For drained loading the expression for Nq for a deep foundation is given by (Vesic
1977)
)]arctan()sin1(tan21[)](tan1[sin1
sin1)tanexp( 2 BLABNq
C8
where A and B are the pile plan dimensions (for a circular pile A = B) and L = pile length For a
circular pile this can be approximated by
asymp 077(120601prime75deg) 119873119902 C9
over a range of effective friction angles 20deg le ϕ le 45deg
C1132 Practical Considerations
For drained loading of sands the full calculated end bearing capacity will never be realized because it
would require the pile to move a distance equal to its diameter This is beyond practical use and
therefore the limit plasticity solutions must be clipped to a fraction of the calculated value Another
reason for using a reduced end-bearing capacity is due to strain incompatibility since the side capacity is
mobilized early while the end-bearing is engaged much later So to have compatible values of Qs and
Qb the qb must be reduced using the following guidelines (Randolph 2003 Mayne 2007)
prime prime C10 119902119887 = 119873119902 ∙ 120590119907119900 ∙ 119891119909
where fx = strain incompatibility factor
fx = 010 for drilled shafts augered cast-in-place and bored piles
fx = 020 for driven low displacement piles (opened-ended steel pipe and H-piles)
fx = 030 for driven high-displacement piles (ie solid piles PSC and closed-ended pipe)
C-12
For undrained loading of circular piles in compression no reduction of end-bearing is necessary
therefore
119902119887 = 933 middot 119904119906 C11
C12 Direct CPT Methods for Pile Capacity
In the direct CPT method the penetrometer readings are scaled directly via specified algorithms to
obtain the pile unit side friction and end-bearing as depicted in Figure C4 As many as 40 different
direct CPT methods have been developed over the past five decades as summarized by Niazi amp Mayne
(2013) Starting circa 1970 many of these early methods relied on hand-recorded data from field
mechanical CPTs where only qc data were obtained at 20 cm intervals or later with mechanical readings
of both qc and fs using special sets of inner and outer rods that recorded vertical load for tip and loads
for tip plus sleeve in alternating increments Also early pile load test data were often obtained from
top-down measurements of load-displacement
C-13
Undrained qb = Nc su
Qside = S (fp DAs)
QTotal = Qs + Qb - Wp
fp = cmck Ko svorsquo tanrsquo
Qbase = qb Ab
Drained qb = Nq svorsquo
Method Two Rational or ldquoIndirectrdquo Method
Method OneldquoDirectrdquo CPT Method
(Scaled Pile)
fp = fctn (soil type pile
type qt fs and u2)
qb = fctn (soil type qt-u2)
qb = unit end bearing
unit sidefriction fp
OCR su Ko t DR rsquo
AXIAL PILE CAPACITY FROM CPT READINGS
Figure C4 Concept of Direct versus Rational Method for CPT evaluation of axial pile capacity
Beginning in the mid-1990s as the modern electric piezocone (CPTu) was implemented the newer
equipment offered improved data and better resolution because of the additional reading of porewater
pressures and correction of raw measured qc to total resistance qt In addition the use of electronic
digital data collection and field computers proved superior in field measurements and recordings As a
consequence several reliable CPT methods for axial pile capacity have been developed Moreover
parallel improvements in full-scale pile load testing occurred and now include modern strain gage
instrumentation digital data recording and automated testing procedures Also the testing can be
single direction (compression or tension) or bi-directional as in the Osterberg cell
C-14
Table C3 provides a selection of recent direct CPT methods that have become available over the
past two decades A number of these (namely ICP NGI UWA and Fugro) were funded by the
offshore industry because of the growth of oil amp gas reserves and windfarm installations thus
necessitating increased concerns on risk probability and reliability in the site investigations for
offshore platforms and design of large driven monopile foundations These direct CPT methods
are often statistically based on large datasets compiled from full-scale load tests made
worldwide (eg Schneider et al 2008)
Table C3 Selection of Direct CPT Methods for Axial Pile Capacity
Method Pile Types Soil Types References CPT
data
Additional
parameters
needed
Unicone Method all types Sands silts
clays
Eslami and
Fellenius
(1997)
Fellenius
(2009)
qt fs u2
KTRI = Kajima
Technical
Research
Institute
all types Sands
mixed clays
Takesue et al
(1998)
fs and u2 No guidance given
on end bearing
resistance
Imperial College
Procedure (ICP)
OE and CE Sands Chow et al
(1997 PhD)
Jardine et al
(2005)
qt Interface friction
angle (δ) from
ring shear tests
Correlated to
mean grain size
(D50)
C-15
OE and CE Clays Jardine et al
(2005)
qt Interface friction
angle (δ)
correlated to PI
Table C3 Continued
NGI Method
(Norwegian
Geotechnical
Institute)
OE and CE Sands Clausen et al
(2005)
qt DR from qt
Clays a Alpha
method
(Karlsrud et al
2005 2012)
b Beta
method
(Karlsrud
2012)
qt for su
qt for
OCR
PI = plasticity
index ()
PI = plasticity
index ()
Fugro Method OE and CE Sands Kolk et al
(2005)
qt
Clays Van Dijk and
Kolk (2011)
qt
OE and CE Sands Lehane et al
(2005)
qt IFR = infilling ratio
related to amount
of plugging
C-16
Purdue LFRD Drilled
shafts
Sands Basu amp
Salgado
(2012)
qt LRFD = load
resistance
factored design
Enhanced Various Sands silts Niazi and qt u2 Based on 330 load
Unicone Method clays and Mayne (2015 and fs tests including
mixed soils 2016) bored augered
jacked and driven
piles
UWA (Univ of
Western
Ra = pile
roughness
Australia) Clays Lehane et al
(2012)
qt Interface friction
angle (δ) from
ring shear tests
and also method
without δ
HKU Method
(Hong Kong
Univ)
OE and CE
(end
resistance
only)
Sands Yu and Yang
(2012)
qt End bearing only
PLR = plug length
ratio
Note PLR can be
estimated from
OE inner diam
Table C3 Continued
Note previously called Marine Technical Directorate (MTD)
C-17
Many of the offshore CPT methods relate to driven piles either in sand or clay as that is
common deep foundation for those purposes Of particular interest are the Unicone and
Modified Unicone Methods since they use all three readings of the piezocone (qt fs and u2)
and address a variety of pile foundation types
C13 Modified UniCone Method
The modified UniCone Method was developed based on a total 330 pile load tests which were
associated with SCPTu data during their site investigations (Niazi and Mayne 2015 2016) This
represents a threefold increase over the original UniCone database that was built upon data
from 106 pile load tests (Eslami amp Fellenius 1997)
For the original UniCone algorithms use is made of the effective cone resistance (qE)
119902119864 = 119902119905 minus 1199062 C12
and a chart of qE vs fs provided an approximate soil classification in five distinct groups as shown by
Figure C5 Later in the modified approach a better delineation of the larger dataset gave soil
subclassifications as indicated by Figure C6
C-18
Figure C5 Soil behavior type using CPT via original UniCone
Figure C6 Soil behavior type using CPT via Modified UniCone
C-19
In the modified approach the 9-zone normalized soil behavioral type (SBTn) is ascertained using CPT
data in conjunction with the Robertson (2009) charts as shown in Figure C7 As discussed in Appendix
A and B the SBTn method uses the normalized cone resistance (Qtn) normalized sleeve friction (Fr())
and CPT material index Ic This permits a much wider range in the calculated pile side friction because fp
is related as a continuous curve with Ic rather than only 5 values that are assigned in the original
scheme
The pile unit side friction (fp) is obtained from qE and the CPT material index Ic using the following
expression at each elevation along the sides of the pile
10(0732 middot 119868119888 minus 3605) fp middot middot C13 = 119902119864 middot 120579119875119879 middot 120579119879119862 120579119877119860119879119864
C-20
Figure C7 Nine-zone soil behavioral soil type using normalized piezocone parameters
(after Robertson 2009 Mayne 2014)
where θPT = coefficient for pile type (θPT = 084 for bored piles 102 for jacked piles 113 for driven
piles) θTC = coefficient for loading direction (θTC = 111 for compression and 085 for tension) and θRATE
= rate coefficient applied to soils in SBT zones 1 through 7 (θRATE = 109 for constant rate of penetration
tests and 097 for maintained load tests) Since CPT provides data at regular intervals of 2 cm to 5 cm
along the sides of the pile the average fp from z = 0 to z = L can be used directly in Equation 1 to obtain
the shaft capacity Qs
C-21
The pile end bearing resistance is obtained from
middot 10(0325middot 119868119888minus1218) 119902119887 = 119902119864 C14
where qE is averaged in the vicinity of the pile tip Figure C8 shows the Modified Unicone Method in
graphical format For sensitive clays of zone 1 please see additional discussions by Niazi amp Mayne
(2016)
Figure C8 Summary of Modified UniCone Method for direct CPT assessment of axial pile
capacity
C-22
C14 Axial Pile Displacement
The movement of pile foundations can be assessed using elastic continuum theory (Poulos amp
Davis 1980 Randolph 2003 Mayne amp Niazi 2017) These relationships have been developed
using finite element analyses boundary elements and analytical closed-form solutions For the
latter approach the top displacement of a rigid pile subjected to a vertical force is shown in
Figure C9 This gives the movement at the top of a pile subjected to either compression or
tension (uplift) loading Also the percentage of axial load transferred to the pile toe is
determined For piles extending through various soil layers the elastic solution can be
implemented by using a set of stacked pile segments each with its own stiffness as
represented by a soil Youngs modulus
For pile groups the use of computer software is recommended Several available programs can handle
pile groups under axial and lateral moment loading such as DEFPIG (Univ Sydney) and PIGLET (Univ
Western Australia) A full listing of pile foundation software is given at the Geotechnical amp
GeoEnvironmental Service Directory wwwggsdcom
Ps
= Sh
aft
Load
sL
t
tEd
IPw
Load Transfer to Base
)]1)((5ln[
)(
)1(1
1
2 vdL
dLFI
E
ECT
211
IF
P
P CT
t
b
Base Load = Pb
Randolph Elastic Solution
Ground Surface
Top Load Pt = Ps + Pb
Rigid Pile Response
wt = displacementd = diameter
L = length
Es = Elastic Soil Modulus
Es0 (at z = 0) surface
EsM (at z = L2)mid-length
EsL (at z = L)full length
Load Transfer to Shaft
E = EsMEsL = Gibson parameterE = 1 for homogeneous case (constant E)E = 05 for pure Gibson case (Es0 = 0)
where FCT = load direction (= 1 compression 0 tension)
21
IF
P
P CT
t
b
z = depth
= Poissons ratio
Tension
Compression
C-23
Figure C9 Elastic continuum solution for axial pile response under compression and tension
loading
C-24
C15 Acknowledgements
The author appreciates the help of Fernando Illingworth of PonsEcuador Meena Viswanth of
GeoSyntecAtlanta and Dr Marco Uzielli of GeoRiskItaly in helping to collect and process the data
Funding support was provided by David Woeller of ConeTec Richmond BC
C-25
- Structure Bookmarks
-
- CONE PENETRATION TEST DESIGN GUIDE FOR STATE GEOTECHNICAL ENGINEERS
- TABLE OF CONTENTS
- LIST OF TABLES
- CHAPTER 1 INTRODUCTION
- CHAPTER 2 DIRECT CPT METHOD FOR SOIL CHARACTERIZATION
- CHAPTER 3 DIRECT CPT METHOD FOR SHALLOW FOUNDATIONS
- CHAPTER 4 DIRECT CPT METHOD FOR DEEP FOUNDATIONS
- REFERENCES
- APPENDIX A
- APPENDIX B
- APPENDIX C
-
To request this document in an alternative format such as braille or large print call 651-366-4718 or 1-800-657-3774 (Greater Minnesota) or email your request to ADArequestdotstatemnus Please request at least one week in advance
Technical Report Documentation Page
1 Report No
MNRC 2018-32
2 3 Recipients Accession No
4 Title and Subtitle
Cone Penetration Test Design Guide for State Geotechnical
Engineers
5 Report Date
November 2018 6
7 Author(s)
Ryan Dagger David Saftner and Paul Mayne
8 Performing Organization Report No
9 Performing Organization Name and Address
Department of Civil Engineering University of Minnesota Duluth 1405 University Drive Duluth MN 55812
10 ProjectTaskWork Unit No
CTS2017022 11 Contract (C) or Grant (G) No
(C) 99008 (WO) 249
12 Sponsoring Organization Name and Address
Minnesota Department of Transportation Research Services amp Library 395 John Ireland Boulevard MS 330 St Paul Minnesota 55155-1899
13 Type of Report and Period Covered
Final Report 14 Sponsoring Agency Code
15 Supplementary Notes
httpmndotgovresearchreports2018201832pdf 16 Abstract (Limit 250 words)
The objectives of this project are focused on a new cone penetration testing (CPT) geotechnical design manual for highway and transportation applications based on recent research and innovation covering the period from 2000 to 2018 A step-by-step procedure is outlined on how to use CPT data in the analysis and design of common geotechnical tasks Previous manuals are either very outdated with information from 1970-1996 or not appropriately targeted to transportation works This design document introduces modern and recent advancements in CPT research not otherwise captured in legacy manuals from the 1990rsquos and earlier Examples and case studies are provided for each topic interpreted using CPT measures In the manual a step-by-step procedure is outlined on how to use CPT data in analysis and design for typical geotechnical practices These topics which are applicable both to state highways and local roads include bridge foundations (including shallow footings and deep foundations) and soil characterization (including determination of standard soil engineering properties)
17 Document AnalysisDescriptors
Cone penetrometers Geographic information systems Soils by
properties Bridge foundations Soil mechanics
18 Availability Statement
No restrictions Document available from
National Technical Information Services
Alexandria Virginia 22312
19 Security Class (this report)
Unclassified
20 Security Class (this page)
Unclassified
21 No of Pages
225
22 Price
CONE PENETRATION TEST DESIGN GUIDE FOR STATE
GEOTECHNICAL ENGINEERS
Prepared by
Ryan Dagger
David Saftner
Department of Civil Engineering
University of Minnesota Duluth
Paul W Mayne
School of Civil amp Environmental Engineering
Georgia Institute of Technology Atlanta
November 2018
Published by
Minnesota Department of Transportation
Research Services amp Library
395 John Ireland Boulevard MS 330
St Paul Minnesota 55155-1899
This report represents the results of research conducted by the authors and does not necessarily represent the views or policies
of the Minnesota Department of Transportation the University of Minnesota Duluth or the Georgia Institute of Technology
This report does not contain a standard or specified technique
The authors the Minnesota Department of Transportation the University of Minnesota Duluth and the Georgia Institute of
Technology do not endorse products or manufacturers Trade or manufacturersrsquo names appear herein solely because they are
considered essential to this report because they are considered essential to this report
TABLE OF CONTENTS
CHAPTER 1 INTRODUCTION1
CHAPTER 2 DIRECT CPT METHOD FOR SOIL CHARACTERIZATION 4
21 Introduction 4
22 Soil Unit Weight 6
23 CPT Material Index 7
231 Step 1 Normalized Sleeve Friction 7
232 Step 2 Iteration 8
24 Soil Behavior Type (SBT) 8
241 Step 1 8
242 Step 2 9
25 Effective Stress Friction Angle 9
26 Stress History 10
27 Lateral Stress Coefficient 11
28 Undrained Shear Strength 12
29 Ground Stiffness and Soil Moduli 12
210 Coefficient of Consolidation 13
211 Hydraulic Conductivity14
212 Example Problems 15
2121 Example 1 Direct CPT Methods for Geoparameters on Sands 15
2122 Example 2 Direct CPT Methods for Geoparameters on Clay 33
CHAPTER 3 DIRECT CPT METHOD FOR SHALLOW FOUNDATIONS 56
31 Procedure 56
311 Step 1 Estimating Footing Dimensions57
312 Step 2 Soil Characterization 58
313 Step 2a Foundation soil formation parameter59
314 Step 3 Soil elastic modulus and Poissonrsquos ratio 60
315 Step 4 Net cone tip resistance 60
316 Step 5 Bearing capacity of the soil 61
317 Step 6 Settlement61
318 Step 7 Final Check 61
32 Example Problems 62
321 Example 3 Direct CPT Method on Sands 62
CHAPTER 4 DIRECT CPT METHOD FOR DEEP FOUNDATIONS 70
41 Introduction70
42 Modified UniCone Method72
421 Step 172
422 Step 272
423 Step 373
424 Step 473
43 Axial Pile Displacements 73
44 Example Problems 74
441 Example 5 Direct CPT Methods Axial Pile Capacity74
REFERENCES 87
APPENDIX A
APPENDIX B
APPENDIX C
1
2
3
4
5
6
7
8
9
10
11
12
13
14
15
16
17
18
19
20
21
22
23
24
LIST OF FIGURES
Figure Geoparameters determined from CPT 1
Figure Conventional method for shallow foundation design compared to direct CPT method 2
Figure Direct CPT evaluation of axial pile capacity 3
Figure Soil unit weight from CPT sleeve friction 6
Figure Approximate ldquorules of thumbrdquo method using CPT sounding from Wakota Bridge MN 7
Figure SBT zones using CPT Ic 9
Figure Yield stress exponent compared to CPT material index 11
Figure k vs dissipation time for 50 consolidation (Mayne 2017) 15
Figure CPT data from Benton County Minnesota for example problem 1 16
Figure Soil layers using ldquorules of thumbrdquo17
Figure Soil layer 1 using SBT method 23
Figure Soil layer 2 using SBT method 24
Figure Soil layer 3 using SBT method 25
Figure Soil layer 4 using SBT method 26
Figure CPT data from Minnesota for example problem 2 34
Figure Soil layers using ldquorules of thumbrdquo36
Figure Soil layer 1 using SBT method 42
Figure Soil layer 2 using SBT method 43
Figure Soil layer 3 using SBT method 44
Figure Soil layer 4 using SBT method 45
Figure Dissipation t50 data 53
Figure Direct CPT method for shallow foundations56
Figure Direct CPT method introduction 58
Figure Differentiation of porewater pressure measurement locations (Lunne et al 1997) 58
Figure 25 Foundation soil formation parameter hs versus CPT material index Ic (Mayne 2017) 60
Figure 26 Diagram of footing profiles for Example 362
Figure 27 CPT data from Northern Minnesota 63
Figure 28 Schematic of foundation design associated with CPT data 64
Figure 29 Soil type for example problem 367
Figure 30 Direct CPT evaluation of axial pile capacity 70
Figure 31 UniCone Method soil behavior type using CPT (Mayne 2017) 71
Figure 32 Modified UniCone Method soil behavior type using CPT (Mayne 2017) 71
Figure 33 Deep foundation end bearing pile diagram74
Figure 34 CPT data from Minnesota for example problem 5 75
Figure 35 Soil layers using ldquorules of thumbrdquo for pile capacity example using direct CPT method 76
Figure 36 Soil layer 1 using SBT method 81
Figure 37 Soil layer 2 using SBT method 82
Figure 38 Soil layer 3 using SBT method 83
Figure 39 Soil layering compared to pilings 85
LIST OF TABLES
Table 1 Geoparameters calculated directly from CPT 4
Table 2 Minor Geoparameters 5
Table 3 Soil type compared to exponent m 10
Table 4 Geoparameters evaluated for Case Example 117
Table 5 Geoparameters 35
Table 6 Bearing capacity defined by soil type61
Table 7 SCPT Results 62
CHAPTER 1 INTRODUCTION
The purpose of this manual is to provide an analysis of soil characterization shallow footings and deep
foundations using direct cone penetration testing (CPT) methods Geotechnical site characterization is
important for evaluating soil parameters that will be used in the analysis and design of foundations
retaining walls embankments and situations involving slope stability Common practice is to determine
these soil parameters through conventional lab and in-situ testing An alternative method uses CPT
readings of cone tip resistance (qt) sleeve friction ( fs) and pore pressure (u2) directly to determine these
parameters such as unit weight effective friction angle undrained shear strength and many others
(Figure 1) Direct CPT methods are provided for these parameters which will be used in the designs for
shallow and deep foundations
Figure 1 Geoparameters determined from CPT
Designing shallow foundations is typically done in a two-part process determining the bearing capacity
and expected settlement (commonly referred to as displacement) of the soil to approximate the
required size and shape of a foundation The older traditional methods are no longer required with
many approaches existing for using CPT directly in the design of shallow foundations Results from these
methods can provide a direct assessment of bearing capacity andor settlement A specific approach to
the direct method has been recommended and tested using a database of 166 full-scale field load tests
(Figure 2) A one-part process is used to scale the measured cone penetrometer readings (ie
measured cone tip resistance sleeve friction and pore water pressure) to obtain the bearing capacity
and settlement of the soil A step-by-step procedure has been created to transition from the CPT data to
bearing capacity with settlement accounted for The steps consist of estimating a design footing width
1
and length while using the process in the Soil Characterization section to determine soil parameters
directly from the CPT data
Figure 2 Conventional method for shallow foundation design compared to direct CPT method
There are upwards of 40 different direct CPT methods that have been developed over the past five
decades to determine the axial compression capacity of a piling foundation Earlier direct CPT methods
relied on hand-recorded information where mechanical-type CPT cone tip resistance data would be
collected at 20 cm intervals
Herein the method recommended for deep foundation design is the Modified UniCone method which
uses all three readings of the modern electronic piezocone penetrometer (CPTu) while addressing a
variety of pile foundation types (Figure 3) The modified UniCone Method is based on a total of 330 pile
load tests (three times the original UniCone database) that were associated with SCPTu data Two
computer software programs have been recommended for analyzing pile movements andor
settlements
2
Figure 3 Direct CPT evaluation of axial pile capacity
3
Symbol Parameter
γt Soil total unit weight
Ic CPT material index
SBT Soil behavior type (SBT)
ʹ σp Preconsolidation stress
YSR Yield stress ratio
ϕʹ Effective friction angle
Ko Lateral stress coefficient
su Undrained shear strength
Dʹ Constrained modulus
Eʹ Drained Youngrsquos modulus
CHAPTER 2 DIRECT CPT METHOD FOR SOIL
CHARACTERIZATION
21 INTRODUCTION
A direct CPT method for determining the value of each of various geoparameters shown in Table 1 is
provided The parameter Ic is used in the derivation of several sequential parameters This section of
the guide may be referred to as these parameters are used in calculations throughout the shallow
foundations and deep foundations sections
Table 1 Geoparameters calculated directly from CPT
4
Kʹ Bulk modulus
ks Subgrade reaction modulus
MR Resilient modulus
Gmax Small-strain shear modulus
k Coefficient of permeability
cv Coefficient of consolidation
Symbol Parameter Equation
ρt Mass Density ρt = γtga
where ga = 98 ms2
σvo Total Stress σvo asymp Σ (γti middot Δzi)
c Effective cohesion ʹ Empirical c asymp 003σp
In clays c asymp 01cu
Other minor parameters can be used in calculations of the geoparameters in Table 1 These minor parameters are provided in Table 2
Table 2 Minor Geoparameters
5
22 SOIL UNIT WEIGHT
The total soil unit weight can be estimated from CPT sleeve friction resistance as shown in Figure 4
(Mayne 2014) This method is not applicable to organic clays diatomaceous soils peats or sensitive
soils
Figure 4 Soil unit weight from CPT sleeve friction
Use Equation 1 to calculate the soils total unit weight
1 120574119905 = 120574119908 ∙ [122 + 015 ∙ ln (100 ∙ 119891119904 + 001)]
120590119886119905119898
Since soil unit weight is required for determining most geoparameters (including Soil Behavior Type)
estimating the soil type of the layers using the ldquorules of thumbrdquo method is a good first step before
determining more precise layering (Figure 5) Once a unit weight is determined for each noticeable layer
change these results can be used in later calculations such as ldquoCPT Material Indexrdquo To use the ldquorules
of thumbrdquo method some helpful guidelines are to assume sands are identified when qt gt 725 psi and u2
asymp uo while the presence of intact clays are prevalent when qt lt 725 psi and u2 gt uo The magnitude of
porewater pressures help to indicate intact clays such as soft (u2 asymp 2middotuo) firm (u2 asymp 4middotuo) stiff (u2 asymp 8middotuo)
and hard (u2 asymp 20middotuo) Fissured overconsolidated clays tend to have negative u2 values such that u2 lt 0
6
Figure 5 Approximate ldquorules of thumbrdquo method using CPT sounding from Wakota Bridge MN
23 CPT MATERIAL INDEX
The development of the CPT material index Ic has improved the initial classification of soil types and
calculation of soil parameters as shown in Table 1 To calculate Ic follow steps 1 and 2
231 Step 1 Normalized Sleeve Friction
Calculate Fr using Equation 2 if not provided with the CPT data gathered
2
7
119891119904 119865119903() = 100 ∙
(119902119905minus120590119907119900)
232 Step 2 Iteration
Iterate using Equations 3-5 to determine Ic by initially using n = 1 to calculate a starting value of Ic The
exponent n is soil-type dependent n = 1 (clays) n asymp 075 (silts) and n asymp 05 (sands) Iteration converges
quickly which is generally after the 3rd cycle
3 (119954119957minus120648119959119952)120648119938119957119950 = 119951 119928119957119951 prime (120648119959119952120648119938119957119950)
4
prime 120590119907119900 119899 = 0381 ∙ 119868119888 + 005 ( ) minus 015 120590119886119905119898
119899 le 10
5 119920119940 = radic[120785 120786120789 minus 119949119952119944119928119957119951]120784 + [120783 120784120784 + 119949119952119944119917119955]120784
24 SOIL BEHAVIOR TYPE (SBT)
Typically soil samples are not taken when using CPT The soil types are then inferred from the qt fs and
u2 readings To determine the types of soil from CPT data follow steps 1 and 2
241 Step 1
To determine the soil layers from CPT results calculate Ic by following the steps under section ldquoCPT
Material Indexrdquo After Ic has been determined through all specified depths use Figure 6 to classify the
type of soil by comparing each Ic value to normalized CPT readings (Fr and Qtn) from Equations 2 and 3
8
Figure 6 SBT zones using CPT Ic
242 Step 2
To determine if any of the soil layers contain ldquosensitive clays and siltsrdquo from zone 1 or ldquovery stiff
overconsolidated (OC) soilrdquo from zones 8 and 9 use Equations 6 and 7 If any soil layers are found within
zone 1 by Equation 6 then caution should be taken as these clays are prone to instability collapses and
difficulties in construction performance Very stiff OC sands to clayey sands of zone 8 (15 lt Fr lt 45)
and very stiff OC clays to silts of zone 9 (Fr gt 45) can be identified by Equation 7
6 Equation 6 errata This is an exponential expression (see Figure 5)
119876119905119899 lt 12119890119909119901(minus14 ∙ 119865119903)
7 119876119905119899 gt 0005(119865119903minus1)minus00003(119865119903minus1)2minus0002
1
Errata See terms and coefficients in Figure 5 above (numbers cannot be rounded off)
After each CPT reading has been assigned a zone from Figure 14 a visual representation can be made to
show the predominant layers by soil types (Figure A5)
25 EFFECTIVE STRESS FRICTION ANGLE
The effective friction angle (ϕ) is used to govern the strength for sands and clays where Equation 8 is
used for sands and Equation 9 is used for clays The value of ϕ for sands is derived from Ic so before ϕ
can be calculated refer to the iteration of Qtn under the ldquoCPT Material Indexrdquo section Once the values
9
Soil Type m
Fissured clays 11
Intact clays 10
Sensitive clays 09
Silt mixtures 085
Silty sands 080
Clean sands 072
Note m may be higher than 11 in fissured clays
of Ic and Qtn are calculated the type of soil can be determined from the section ldquoSoil Behavior Type
SBTrdquo The type of soil will dictate which equation to use for ϕ
Sands
8 120601prime(deg) = 176deg + 110deg log(119876119905119899)
Clays
9 119902119905minus120590119907119900 120601prime(deg) = 295deg ∙ 119861119902
0121 ∙ [0256 + 0336 ∙ 119861119902 + log ( )] prime 120590119907119900
Where = 119861119902 (1199062 minus 119906119900)(119902119905 minus 120590119907119900)
26 STRESS HISTORY
Determining the stress history can be characterized by an apparent yield stress ratio of the form
10 prime 120590119901
119884119878119877 = prime 120590119907119900
Where σp is defined as the preconsolidation stress or effective yield stress (Equation 10) The YSR is the
same equation as the more common overconsolidation ratio (OCR) but is now generalized to
accommodate mechanisms of preconsolidation such as ageing desiccation repeated cycles of wetting-
drying repeated freeze thaw cycles and other factors
11 prime prime 120590119901 = 120590119910 = 033(119902119905 minus 120590119907119900)119898prime
Where σp σy σvo and qt have units of kPa and the value of m depends on soil type with typical values shown in Table 3
Table 3 Soil type compared to exponent m
10
The value of m for non-fissured soils and inorganic clays and silts is derived from Ic (Figure 7) so before
m can be calculated (Equation 12) refer to the iteration of Ic under the ldquoCPT Material Indexrdquo section
Determine Ic for all soil layers before calculating m
12 028
119898prime = 1 minus 25 119868119888 1+( frasl ) 265
Once m is known for all soil layers the YSR can be determined
Figure 7 Yield stress exponent compared to CPT material index
A limiting value of YSR can be reached for clays and sands It can be calculated using Equation 13
13 (1 sin(emptyprime))
(1+sin(emptyprime)) 119884119878119877119897119894119898119894119905 = [ ]
(1minussin(emptyprime))2
27 LATERAL STRESS COEFFICIENT
The lateral stress coefficient Ko = σhoσvo commonly referred to as the at-rest condition is used to
represent the horizontal geostatic state of soil stress Ko can be calculated using Equation 14 but
11
14
119870119900 = (1 minus sin(emptyprime)) ∙ 119884119878119877sin( emptyprime)
sections such as ldquoStress Historyrdquo and ldquoEffective Stress Friction Anglerdquo will need to be referred to for
determining parameters ϕ and YSR
A maximum value for Ko can be determined by Equation 15
15 (1+sin(emptyprime))
119870119900119898119886119909 = = 1199051198861198992(45deg + emptyprime2) (1minussin(emptyprime))
28 UNDRAINED SHEAR STRENGTH
Loading on soils can result in fully drained partially drained or fully undrained conditions Sands
typically produce drained cases due to their high permeability but exceptions may occur in loose sands
during fast loading where the water does not have sufficient time to dissipate Clays exhibit low
permeability and thus often result in undrained loading cases when a load is applied quickly For soft-
firm clays the undrained shear strength (su) can be determined from CPT via Equation 16 where the
value of the bearing factor Nkt can be taken as 12
16 119902119905minus120590119907119900 119904119906 =
119873119896119905
In the case of remolded undrained shear strength from CPT su asymp fs
29 GROUND STIFFNESS AND SOIL MODULI
Determining the grounds stiffness can be measured from geoparameters such as the constrained
modulus (D) drained Youngrsquos modulus (E) bulk modulus (K) subgrade reaction modulus (ks) resilient
modulus (MR) and small-strain shear modulus (Gmax)
17 119863 prime asymp 5 ∙ (119902119905 minus 120590119907119900)
18
12
119864 prime = 119863 prime
11
19 119864prime
119870prime = [3∙(1minus2119907prime)]
The subgrade modulus (ks) is a combination of soil-structural properties which creates a parameter that
depends on the ground stiffness and the size of the loaded element
20 119864prime
119896119904 = [119889∙(1minus1199072)]
The resilient modulus MR applies to pavement analysis and design and can be calculated using Equation
21 where qt and fs are in MPa
21 053 119872119877 = (146119902119905 + 135511989111990414 + 236)244
The small strain shear modulus Gmax is a representation of the initial stiffness of all soils and rocks
Graphically it is the beginning portion of all stress-strain-strength curves for geomaterials Use Equation
22 to determine Gmax
22 119866119898119886119909 = 120588119905 ∙ 1198811199042
Where Vs = shear wave velocity (Equation 23) as measured by seismic cone penetration tests (SCPT) If
only cone penetration tests (CPT) or piezocone (CPTu) data are available the shear wave velocity may
be estimated from
23 03 119891119904 119881119904 (119898frasl119904) = [101 ∙ log(119902119905) minus 114]167 ∙ (100 ∙ )
119902119905
Where qt and fs have units of (kPa)
210 COEFFICIENT OF CONSOLIDATION
The coefficient of consolidation (cv) controls the rate that foundation and embankment settlements
occur By using results of CPT dissipation tests that measure the change in u2 readings over time the
value of cv can be determined Using Equation 24 cv can be determined based on piezocone dissipation
13
curves The equation below requires an estimate of the in-situ rigidity index (IR) of the soil If results of
SCPTU are available then IR may be determined from Gmax and qt per equation A38
24 0030∙(119886119888)2∙(119868119877)075
119888119907 = 11990550
Where ac = penetrometer radius (178 cm for 10-cm2 cone 220 cm for 15-cm2 cone) t50 = time to reach 50 dissipation IR = Gsu = undrained rigidity index G can be determined from the ldquoGround Stiffness and Soil Modulirdquo section
211 HYDRAULIC CONDUCTIVITY
The hydraulic conductivity also known as the coefficient of permeability (k) expresses the flow
characteristics of soils and has units of cms or feetday One method of calculation would be to use
Equation 25 where cv would need to be determined from ldquoCoefficient of Consolidationrdquo and D would
need to be determined from ldquoGround Stiffness and Soil Modulirdquo
25 119888119907∙120574119908 119896 =
119863prime
An alternative approach developed for soft normally-consolidated soils (Figure 8) is shown in Equation
26 where t50 (sec) values are used directly in assessing k in (cms)
26
14
125
119896 asymp ( 1
) 251∙11990550
Figure 8 k vs dissipation time for 50 consolidation (Mayne 2017)
212 EXAMPLE PROBLEMS
2121 Example 1 Direct CPT Methods for Geoparameters on Sands
Several geoparameters need to be determined based on the given CPT data collected for the South
Abutment of a bridge in Benton County MN (Figure 9) The groundwater table (GWT) was measured at
17 feet Determine all the geoparameters found in Table 4 at depths of 0 feet to 30 feet All sand layers
can be assumed ldquodrainedrdquo with ν = 02
15
Figure 9 CPT data from Benton County Minnesota for example problem 1
16
Table 4 Geoparameters evaluated for Case Example 1
Symbol Parameter Symbol Parameter
γt Soil total unit weight Ko Lateral stress coefficient
Ic CPT material index Dʹ Constrained modulus
SBT Soil behavior type (SBT) Eʹ Drained Youngrsquos modulus ʹ σp Preconsolidation stress Kʹ Bulk modulus
YSR Yield stress ratio MR Resilient modulus
ϕʹ Effective friction angle Gmax Small-strain shear modulus
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 10 Soil layers using ldquorules of thumbrdquo
γw = 6224 pcf
17
Layer 1
From Figure 10 fs = 17 psi taken as a representative value of the layer
17 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1204 pcf
145 psi
Layer 2
From Figure 10 fs = 17 psi taken as a representative value of the layer
17psi
γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1204 pcf 145 psi
Layer 3
From Figure 10 fs = 12 psi taken as a representative value of the layer
12 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 4
From Figure 10 fs = 7 psi taken as a representative value of the layer
CPT Material Index
7 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1121 pcf
145 psi
Layer 1
From Figure 10 fs = 17 psi and qc = 3500 psi
18
qt = qc + u2 ∙ (1 minus a) = 3500 psi + 3 psi ∙ (1 minus 08) = 3501 psi
= γt ∙ 6 feet = 1204 pcf ∙ 6 feet = 7225 psf = 50 psi σvo
fs 17 psi Fr() = 100 ∙ = 100 ∙ = 049
(qt minus σvo) (3501 psi minus 50 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 5 psi minus 0 psi = 50 psi
(qt minus σvo)σatm (3501 psi minus 50 psi )145 psi = = Qtn prime )n (σvoσatm (50 psi145 psi)n
prime σvo 50 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(049)]2
Qtn = 35329 n = 036 lt 10 Ic = 13
Layer 2
From Figure 10 fs = 17 psi and qc = 3500 psi
19
qt = qc + u2 ∙ (1 minus a) = 3500 psi + 0 psi ∙ (1 minus 08) = 3500 psi
= 7225 psf + γt ∙ 6 feet = 7225 psf + 1204 pcf ∙ 6 feet = 1445 psf = 100 psi σvo
fs 17 psi Fr() = 100 ∙ = 100 ∙ = 049
(qt minus σvo) (3500 psi minus 100 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 100 psi minus 0 psi = 100 psi
(qt minus σvo)σatm (3500 psi minus 100 psi )145 psi = = Qtn prime )n (σvoσatm (100 psi145 psi)n
prime σvo 100 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
47 minus log(Qtn)]2 + [122 + log(049)]2
Ic = radic[3
Qtn = 27947 n = 041 lt 10 Ic = 14
Layer 3
From Figure 10 fs = 12 psi and qc = 1500 psi
20
qt = qc + u2 ∙ (1 minus a) = 1500 psi + 0 psi ∙ (1 minus 08) = 1500 psi
= 1445 psf + γt ∙ 11 feet = 1445 psf + 1172 pcf ∙ 11 feet = 2734 psf = 190 psi σvo
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 081
(qt minus σvo) (1500 psi minus 190 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = γwater ∙ (z minus 119911119908) = 6224 pcf ∙ (23 feet minus 17 feet) = 3734 psf = 99 psi
prime σvo = σvo minus uo = 190 psi minus 99 psi = 164 psi
(qt minus σvo)σatm (1500 psi minus 190 psi )145 psi = = Qtn prime )n (σvoσatm (164 psi145 psi)n
prime σvo 164 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(081)]2
Qtn = 947 n = 06 lt 10 Ic = 19
21
Layer 4
From Figure 10 fs = 7 psi and qc = 1200 psi
qt = qc + u2 ∙ (1 minus a) = 1200 psi + 0 psi ∙ (1 minus 08) = 1200 psi
= 2734 psf + γt ∙ 11 feet = 2734 psf + 1121 pcf ∙ 7 feet = 3519 psf = 244 psi σvo
fs 7 psi Fr() = 100 ∙ = 100 ∙ = 060
(qt minus σvo) (1200 psi minus 244 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = γwater ∙ (z minus 119911119908) = 6224 pcf ∙ (30 feet minus 17 feet) = 8091 psf = 130 psi
prime σvo = σvo minus uo = 244 psi minus 130 psi = 188 psi
(qt minus σvo)σatm (1200 psi minus 244 psi )145 psi = = Qtn prime )n (σvoσatm (188 psi145 psi)n
prime σvo 188 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(060)]2
22
Qtn = 686 n = 064 lt 10 Ic = 19
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic Qtn and Fr the first layer is defined as a ldquoDrained Gravelly Sandrdquo from
Figure 11
Figure 11 Soil layer 1 using SBT method
23
Layer 2
Based on values of Ic Qtn and Fr the second layer is defined as a ldquoDrained Sandrdquo from Figure
12
Figure 12 Soil layer 2 using SBT method
24
Layer 3
Based on values of Ic Qtn and Fr the third layer is defined as a ldquoDrained Sandrdquo from Figure 13
Figure 13 Soil layer 3 using SBT method
25
Layer 4
Based on values of Ic Qtn and Fr the fourth layer is defined as a ldquoDrained Sandrdquo from Figure 14
Figure 14 Soil layer 4 using SBT method
Effective Stress Friction Angle
Layer 1
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(3533) = 456deg
Layer 2
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(2795) = 445deg
26
Layer 3
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(947) = 393deg
Layer 4
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(686) = 378deg
Stress History
Layer 1
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (13frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(24139 kPa minus 345 kPa)072 = 4717 kPa
= 684 psi
prime σp 684 psi YSR = = = 136
primeσvo 50 psi
Layer 2
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + (14 1 + ( frasl ) frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(34132 kPa minus 689 kPa)072 = 605 kPa
= 877 psi
prime σp 877 psi YSR = = = 87
primeσvo 100 psi
27
Layer 3
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + (19 1 + ( frasl ) frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(10342 kPa minus 131 kPa)072 = 254 kPa
= 368 psi
prime σp 368 psi YSR = = = 22
primeσvo 164 psi
Layer 4
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (19frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(8274 kPa minus 168 kPa)072 = 215 kPa = 312 psi
prime σp 312 psi YSR = = = 17
primeσvo 188 psi
Lateral Stress Coefficient
Layer 1
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(456deg)) ∙ 136sin( 456deg) = 18
Layer 2
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(445deg)) ∙ 87sin( 445deg) = 14
28
Dprime 17480 psi
Eprime = = = 15890 psi 11 11
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(393deg)) ∙ 22sin( 393deg) = 06
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(378deg)) ∙ 17sin( 378deg) = 05
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3501 psi minus 50 psi) = 17480 psi
Eprime 15890 psi
Kprime = = = 8828 psi [3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Layer 3
Layer 4
Ground Stiffness and Soil Moduli
Layer 1
Values of qt and fs need to be in MPa
MR = (146(241 MPa)053 + 1355(012 MPa)14 + 236)244 = 3416 MPa = 49575 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1172 kPa m Vs (mfrasls) = [101 ∙ log(24139 kPa) minus 114]167 ∙ (100 ∙ ) = 2745
24139 kPa s
Vs = 9012 fts
29
γt 1204 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 37 ∙ (9012 ) = 304 ∙ 106psf = 21000 psi s
Layer 2
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3500 psi minus 100 psi) = 17480 psi
Dprime 17480 psi Eprime = = = 15890 psi
11 11
Eprime 15890 psi Kprime = = = 8828 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(241 MPa)053 + 1355(012 MPa)14 + 236)244 = 3415 MPa = 49562 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1172 kPa m Vs (mfrasls) = [101 ∙ log(24133 kPa) minus 114]167 ∙ (100 ∙ ) = 2745
24133 kPa s
Vs = 9012 fts
γt 1204 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
30
2 ft Gmax = ρt ∙ Vs
2 = 37 ∙ (9012 ) = 304 ∙ 106psf = 21000psi s
Layer 3
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (1500 psi minus 190 psi) = 7405 psi
Dprime 7405 psi Eprime = = = 6732 psi
11 11
Eprime 6732 psi Kprime = = = 3740 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(103 MPa)053 + 1355(008 MPa)14 + 236)244 = 1506 MPa = 21854 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 827 kPa m Vs (mfrasls) = [101 ∙ log(10343 kPa) minus 114]167 ∙ (100 ∙ ) = 2611
10343 kPa s
Vs = 8566 fts
γt 1172 pcf slug ρt = = = 36
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 36 ∙ (8566 ) = 27 ∙ 106psf = 19000 psi s
Layer 4
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (1200 psi minus 244 psi) = 5878 psi
31
Dprime 5878 psi Eprime = = = 5343 psi
11 11
Eprime 5343 psi Kprime = = = 2969 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(83 MPa)053 + 1355(005 MPa)14 + 236)244 = 165 MPa = 16901 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 483 kPa m Vs (mfrasls) = [101 ∙ log(8274 kPa) minus 114]167 ∙ (100 ∙ ) = 2243
8274 kPa s
Vs = 7360 fts
γt 1121 pcf slug ρt = = = 35
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 35 ∙ (7360 ) = 389 ∙ 106psf = 13000 psi s
32
2122 Example 2 Direct CPT Methods for Geoparameters on Clay
Several geoparameters need to be determined based on the given CPT data collected for the South
Abutment (Figure 15) The groundwater table (GWT) was measured at 60 feet Determine all the
geoparameters found in Table 5 at depths of 0 feet to 42 feet All sand layers can be assumed ldquodrainedrdquo
ν = 02 All clay layers can assume ν = 049
33
Figure 15 CPT data from Minnesota for example problem 2
34
Table 5 Geoparameters
Symbol Parameter Symbol Parameter
γt Soil total unit weight Eʹ Drained Youngrsquos modulus
Ic CPT material index Kʹ Bulk modulus
SBT Soil behavior type (SBT) MR Resilient modulus
ʹ σp Preconsolidation stress Gmax Small-strain shear modulus
YSR Yield stress ratio su Undrained shear strength
ϕʹ Effective friction angle cv Coefficient of consolidation
Ko Lateral stress coefficient k Hydraulic conductivity
Dʹ Constrained modulus
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
35
Figure 16 Soil layers using ldquorules of thumbrdquo
Unit weight of water γw = 6224 pcf
Layer 1
From Figure 16 fs = 13 psi taken as a representative value of the layer
13 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1179 pcf
145 psi
Layer 2
From Figure 16 fs = 12 psi taken as a representative value of the layer
12 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 3
From Figure 16 fs = 20 psi taken as a representative value of the layer
36
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
Layer 4
From Figure 16 fs = 2 psi taken as a representative value of the layer
2 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1004 pcf
145 psi
CPT Material Index
Layer 1
From Figure 10 fs = 13 psi and qc = 3000 psi
qt = qc + u2 ∙ (1 minus a) = 3000 psi + 0 psi ∙ (1 minus 08) = 3000 psi
σvo = γt ∙ 2 feet = 1179 pcf ∙ 2 feet = 235 psf = 16 psi
fs 13 psi Fr() = 100 ∙ = 100 ∙ = 043
(qt minus σvo) (3000 psi minus 16 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 16 psi minus 0 psi = 16 psi
(qt minus σvo)σatm (3000 psi minus 16 psi )145 psi = = Qtn prime )n (16 psi145 psi)n (σvoσatm
prime σvo 16 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(043)]2
37
Qtn = 4126 n = 032 lt 10 Ic = 121 therefore sand
Layer 2
From Figure 10 fs = 12 psi and qc = 250 psi
qt = qc + u2 ∙ (1 minus a) = 250 psi + 10 psi ∙ (1 minus 08) = 252 psi
σvo = 235 psf + γt ∙ 30 feet = 235 psf + 1172 pcf ∙ 30 feet = 3751 psf = 260 psi
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 53
(q minus σ ) (252 psi minus 260 psi) t vo
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 260 psi minus 0 psi = 260 psi
(qt minus σvo)σatm (250 psi minus 260 psi )145 psi = = Qtn prime (σvoσatm)n (260 psi145 psi)n
38
prime σvo 260 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(53)]2
Qtn = 87 n = 10 le 10 Ic = 32 therefore clay
Layer 3
From Figure 10 fs = 20 psi and qc = 4000 psi
qt = qc + u2 ∙ (1 minus a) = 4000 psi + 8 psi ∙ (1 minus 08) = 40016 psi
σvo = 3751 psf + γt ∙ 2 feet = 3751 psf + 1219 pcf ∙ 2 feet = 3995 psf = 277 psi
fs 16 psi Fr() = 100 ∙ = 100 ∙ = 054
(qt minus σvo) (30016 psi minus 277 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 277 psi minus 0 psi = 277 psi
(qt minus σvo)σatm (30016 minus 277)145 psi = = Qtn prime (σvoσatm)n (277 psi145 psi)n
39
primeσvo 277 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(054)]2
Qtn = 1962 n = 052 le 10 Ic = 15 therefore sand
Layer 4
From Figure 10 fs = 2 psi and qc = 250 psi
qt = qc + u2 ∙ (1 minus a) = 250 psi + 0 psi ∙ (1 minus 08) = 250 psi
= 3994 psf + γt ∙ 8 feet = 3994 psf + 1004 pcf ∙ 8 feet = 4798 psf = 333 psi σvo
fs 2 psi Fr() = 100 ∙ = 100 ∙ = 092
(qt minus σvo) (250 psi minus 333 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 333 psi minus 0 psi = 333 psi
(qt minus σvo)σatm (250 psi minus 333 psi )145 psi = = Qtn prime (σvoσatm)n (333 psi145 psi)n
40
prime σvo 333 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(092)]2
Qtn = 65 n = 10 lt 10 Ic = 29 therefore clayey silt
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic Qtn and Fr the first layer is defined as a ldquoDrained Sandrdquo from Figure 17
41
Figure 17 Soil layer 1 using SBT method
Layer 2
Based on values of Ic Qtn and Fr the second layer is defined as a ldquoUndrained Clayrdquo from Figure
18
42
Figure 18 Soil layer 2 using SBT method
43
Layer 3
Based on values of Ic Qtn and Fr the third layer is defined as a ldquoDrained Sandrdquo from Figure 19
Figure 19 Soil layer 3 using SBT method
44
Layer 4
Based on values of Ic Qtn and Fr the fourth layer is defined as a ldquoUndrained Silty Mixrdquo from
Figure 20
Figure 20 Soil layer 4 using SBT method
Effective Stress Friction Angle
Layer 1 (sand)
45
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(4126) = 464deg
Layer 2 (clay)
qt minus σvo
ϕprime(deg) = 295deg ∙ Bq0121 ∙ [0256 + 0336 ∙ Bq + log ( )]
primeσvo
Where
(u2 minus uo) (10 psi minus 0 psi) Bq = = = 004
(qt minus σvo) (252 psi minus 260 psi)
252 psi minus 260 psi ϕprime(deg) = 295deg ∙ 00440121 ∙ [0256 + 0336 ∙ 0044 + log ( )]
260 psi
ϕprime(deg) = 245deg
Layer 3 (sand)
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(1962) = 428deg
Layer 4 (clay)
qt minus σvo
ϕprime(deg) = 295deg ∙ Bq0121 ∙ [0256 + 0336 ∙ Bq + log ( )]
primeσvo
Where
(u2 minus uo) (5 psi minus 0 psi) Bq = = = 002
(qt minus σvo) (251 psi minus 333 psi)
252 psi minus 333 psi ϕprime(deg) = 295deg ∙ 0020121 ∙ [0256 + 0336 ∙ 002 + log ( )]
333 psi
ϕprime(deg) = 265deg
Stress History
46
Layer 1
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (12frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(3000 psi minus 33)072 = 1051 psi
prime σp 1051 psi YSR = = = 642
primeσvo 16 psi
Layer 2
028 028
prime m = 1 minus = 1 minus = 099 25 25 1 + (
Icfrasl ) 1 + (32frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(252 psi minus 260)099 = 734 psi
prime σp 734 psi YSR = = = 28
primeσvo 260 psi
Layer 3
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + ( frasl ) 1 + (15frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(4002 psi minus 277)072 = 1288 psi
47
prime σp 1288 psi YSR = = = 46
primeσvo 277 psi
Layer 4
028 028
prime m = 1 minus = 1 minus = 097 25 25 Ic 1 + ( frasl ) 1 + (29frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(250 psi minus 333)097 = 626 psi
prime σp 626 psi YSR = = = 19
primeσvo 333 psi
Lateral Stress Coefficient
Layer 1
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(464deg)) ∙ 642sin( 464deg) = 56
Layer 2
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(245deg)) ∙ 28sin( 245deg) = 090
Layer 3
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(428deg)) ∙ 46sin( 428deg) = 091
Layer 4
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(266deg)) ∙ 19sin( 266deg) = 073
48
Ground Stiffness and Soil Moduli
Layer 1
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3000 psi minus 16 psi) = 14991 psi
Dprime 14991 psi Eprime = = = 13629 psi
11 11
Eprime 13629 psi Kprime = = = 7571 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(207 MPa)053 + 1355(009 MPa)14 + 236)244 = 2816 MPa = 40873 psi
fs
03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 896 kPa m Vs (mfrasls) = [101 ∙ log(20685 kPa) minus 114]167 ∙ (100 ∙ ) = 2564
20685 kPa s
Vs = 8412 fts
γt 1179 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
2 ft 2 Gmax = ρt ∙ Vs = 37 ∙ (8412 ) = 260 ∙ 106psf = 17992 psi
s
Layer 2
49
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (252 psi minus 260 psi) = 11298 psi
Dprime 11298 psi Eprime = = = 10271 psi
11 11
Eprime 10271 psi Kprime = = = 17118 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(049))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(17 MPa)053 + 1355(008 MPa)14 + 236)244 = 443 MPa = 64309 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 827 kPa m Vs (mfrasls) = [101 ∙ log(17375 kPa) minus 114]167 ∙ (100 ∙ ) = 2646
17375 kPa s
Vs = 8680 fts
γt 1172 pcf slug ρt = = = 36
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 36 ∙ (8680 ) = 274 ∙ 106psf = 19036 psi s
Layer 3
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (4002 psi minus 277 psi) = 19869 psi
Dprime 19869 psi Eprime = = = 18063 psi
11 11
50
Eprime 18063 psi Kprime = = = 10035 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(276 MPa)053 + 1355(014 MPa)14 + 236)244 = 4017 MPa = 58309 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1379 kPa m Vs (mfrasls) = [101 ∙ log(27591 kPa) minus 1379]167 ∙ (100 ∙ ) = 2854
27591 kPa s
Vs = 9363 fts
γt 121 pcf slug ρt = = = 38
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 38 ∙ (9363 ) = 33 ∙ 106psf = 23050 psi s
Layer 4
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (250 psi minus 333 psi) = 1088 psi
Dprime 1088 psi Eprime = = = 989 psi
11 11
Eprime 989 psi Kprime = = = 16490 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(049))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
51
MR = (146(17 MPa)053 + 1355(001 MPa)14 + 236)244 = 360 MPa = 52189 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 138 kPa m Vs (mfrasls) = [101 ∙ log(17238 kPa) minus 114]167 ∙ (100 ∙ ) = 1545
17238 kPa s
Vs = 5069 fts
γt 1004 pcf slug ρt = = = 31
ga 322 ftfrasls2 ft3
2 ft 2 Gmax = ρt ∙ Vs = 31 ∙ (5069 ) = 80 ∙ 105psf = 5565 psi
s
Undrained Shear Strength
Layer 2
qt minus σv0 250 psi minus 26 psi su = = = 187 psi
12 12
Layer 4
qt minus σv0 250 psi minus 333 psi su = = = 181 psi
12 12
Coefficient of Consolidation
52
Example dissipation data are shown in Figure 21 Example calculations are provided
Figure 21 Dissipation t50 data
Layer 1
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (7197)075
cv = = = 359 cmsec 1 sec t50
Layer 2
53
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (94392)075
cv = = = 0008 cmsec 3000 sec t50
Layer 3
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (8303)075
cv = = = 665 cmsec 06 sec t50
Layer 4
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (26714)075
cv = = = 087 cmsec 11 sec t50
Hydraulic Conductivity
Layer 1
cm 1 psi 359 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 10 ∙ 10minus4 cmsec Dprime 14984 psi
Layer 2
cm 1 psi 0008 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 73 ∙ 10minus7 cmsec Dprime 11379 psi
Layer 3
cm 1 psi 665 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 20 ∙ 10minus5 cmsec Dprime 148650 psi
54
Layer 4
cm 1 psi 087 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 18 ∙ 10minus4 cmsec Dprime 10861 psi
55
CHAPTER 3 DIRECT CPT METHOD FOR SHALLOW
FOUNDATIONS
31 PROCEDURE
Shallow foundation analysis is typically done in a two-part traditional procedure The traditional
techniques are no longer required as a direct CPT method for square rectangular and circular shallow
footings is available (Figure 22) This process has the soil types grouped into four main categories sands
silts fissured clays and intact clays When determining soil types for each design it is believed footings
on sands and silts act in a fully drained manner while intact clays act in an undrained manner under
conditions of constant volume In order to determine the vertical stress-displacement-capacity of
square rectangular and circular shallow footings follow the steps provided towards the solution given
by Equation 27
Figure 22 Direct CPT method for shallow foundations
Equation 27 may be used to calculate all footing stresses from zero to the bearing capacity (qmax) To
calculate qmax for a sized footing of width (B) length (L) and thickness (t) follow steps 1 through 7
provided The settlement (s) can be determined after the calculation of qmax by simply rearranging
56
Equation 27 where the allowable stress (qallow) is defined as qmax divided by the factor of safety (FS) For
shallow footings a FS value of 3 is commonly used in geotechnical engineering
120782120787 minus120782120785120786120787 119956 119923 119954119950119938119961 = 119945119956 ∙ 119954119957119951119942119957 ∙ (119913
) ∙ (119913
) 27 119950119938119961
2 119902119898119886119909 0345 1 frasl119865119878 119871 119904 = 119861 ∙ [ ∙ ∙ ( ) ] 28
119861 ℎ119904 119902119905119899119890119905
Where hs = the foundation soil formation parameter qtnet = the net corrected cone tip resistance
311 Step 1 Estimating Footing Dimensions
In Minnesota frost heave can have devastating effect on a shallow foundations It is common practice to
place a foundation bearing elevation below the expected maximum frost depth (roughly 45 to 6 feet
below ground level) With this assumption estimate a footing size (B x L) for design to obtain
representative data from CPT roughly 15middotB below the foundation depth (Df) as shown in Figure 23 The
CPT data collected will consist of
cone tip resistance (qc)
measured porewater pressure acting behind the cone tip (u2) as shown in Figure 24
sleeve friction (fs)
57
Figure 23 Direct CPT method introduction
Figure 24 Differentiation of porewater pressure measurement locations (Lunne et al 1997)
312 Step 2 Soil Characterization
The soil behavior type (SBT) and CPT material index (Ic) govern the value of the formation factor hs The
first step is following the steps in Soil Unit Weight to determine soil layering and γt values for each
layer To determine a representative unit weight (γsoil) for all the layers in the range of Df to 15∙B below
58
Df the user will need to use their engineering judgement on the definition of ldquorepresentative unit
weightrdquo This single value of unit weight is used in further calculations such as the total vertical soil stress (σvo) from Equation 29 and effective vertical stress (σvo) from Equation 30 Values of σvo and σvo
will need to be calculated at 15middotB below Df
120648119959119952 = sum(120632119956119952119946119949 ∙ 119963) 29
prime 120648119959119952 = 120648119959119952 minus 119958119952 30
Where z = Df +15middotB uo=γwatermiddot(z-zw)
The qc will need to be corrected using Equation 31 in the case of fine-grained soils that develop excess porewater pressure during cone penetration These values will be used in the following calculations
119954119957 = 119954119940 + 119958120784 ∙ (120783 minus 119938) 31
Where 119860119899 a = cone area ratio = eg MnDOT commonly uses a = 08 119860119888
119860119899 = cross-sectional area of load cell or shaft 119860119888 = projected area of the cone
The cone area ratio is determined based on the type of piezocone tip used during in-situ field testing
Manufacturer specifications should provide the measured net area ratio (a) for the particular cone
penetrometer as determined by calibration in a pressurized triaxial chamber
313 Step 2a Foundation soil formation parameter
With the representative value γsoil continue to follow the steps in CPT Material Index through Soil
Behavior Type (SBT) to better define the type of soil at depth 15∙B below Df Use the value of Ic
calculated at depth 15∙B below Df to calculate hs with Equation 36 This parameter is based on the soil
type with typical values shown in Figure 25 The data on silts and sands are considered fully drained
whereas the fissured clay subset may be partially drained to undrained
59
23 ℎ119904 = 28 minus
119868119888 15 32
1+( ) 24
Figure 25 Foundation soil formation parameter hs versus CPT material index Ic (Mayne 2017)
314 Step 3 Soil elastic modulus and Poissonrsquos ratio
Details about the soil such as its elastic modulus (Es) and Poissonrsquos ratio (ν) will be needed for further
calculations A representative value for the Es can be determined from in-situ field tests Values of ν can
be assigned as 02 for drained sands and as 05 for undrained loading cases involving clays (Jardine et al
1985 Burland 1989)
315 Step 4 Net cone tip resistance
Calculate qtnet the mean value of net cone tip resistance 15middotB below the foundation bearing elevation
using Equation 33
33 119902119905119899119890119905 = 119902119905 minus 120590119907119900
60
316 Step 5 Bearing capacity of the soil
Use the assumed B and L calculated hs and qtnet to determine the soils bearing capacity (qmax) from
Equation 34 Use Table 6 to calculate qmax by using the maximum allowable settlement ratio (sB)max
correlating to the soil type If hs is in between the given values interpolate to acquire (sB)max
Table 6 Bearing capacity defined by soil type
Type of Soil hs (sB)max
Clean Sands 058 12
Silts 112 10
Fissured Clays 147 7
Intact Clays 270 4
05 minus0345 119904 119871 119902119898119886119909 = ℎ119904 ∙ 119902119905119899119890119905 ∙ (119861
) ∙ (119861
) 34 119898119886119909
317 Step 6 Settlement
Settlement can be calculated directly using the results from Equation 35 A FS equal to 3 is common in
foundation engineering
2 119902119898119886119909 0345 1 frasl119865119878 119871 119904 = 119861 ∙ [ ∙ ∙ ( ) ] 35
119861 ℎ119904 119902119905119899119890119905
318 Step 7 Final Check
Check that the applied stress (q) is less than qmax using Equation 36 Repeat the process again with a
new B andor new L if q gt qmax
05 119904 ) 119902 = ℎ119904 ∙ 119902119905119899119890119905 ∙ (119861
minus0345 119871 ∙ ( )
11986136
61
32 EXAMPLE PROBLEMS
321 Example 3 Direct CPT Method on Sands
A footing size needs to be determined based on the given CPT data collected for the South Abutment
(Figure 27) The footing stress (q) was determined to be 8000 psf Estimate a footing size (B x L) and
determine the bearing capacity of the foundation (qmax) using the direct CPT method provided Also
determine the expected settlement based on the calculated bearing capacity
Figure 26 Diagram of footing profiles for Example 3
The soil elastic modulus determined from seismic CPT (SCPT) are shown in Table 7
Table 7 SCPT Results
Depth (feet) Es (tsf)
Bottom of layer
3 557
6 433
9 557
12 695
17 590
22 501
27 571
32 505
62
Figure 27 CPT data from Northern Minnesota
Solution
Step 1 Assume the footing is placed below the frost depth of 6 feet
Estimate L L = 50 feet = 600 inches
63
Estimate B B = 12 feet = 144 inches
Estimate footing thickness t t = 2 feet
Df = 6 feet = 72 inches
Df + 15middotB = 24 feet = 288 inches
Step 2 Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 28 Schematic of foundation design associated with CPT data
Layer 1
From Figure 28 fs = 20 psi taken as a representative value of the layer
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
64
Layer 2
From Figure 28 fs = 20 psi taken as a representative value of the layer
20psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
Layer 3
From Figure 28 fs = 8 psi taken as a representative value of the layer
8 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1134 pcf
145 psi
Using the unit weights calculated between Df and 15B determine a representative unit weight of the soil to calculate the total and effective soil stresses Based on layer 3 being the weakest supporting layer γsoil = 113 pcf
σvo = γsoil ∙ (119863119891 + 15 ∙ B) = 113 lbsfraslft3 ∙ 24 feet = 2721 psf = 189 psi
prime σvo = σvo minus uo = 2721 psf minus 0 psf = 2721 psf = 189 psi
Calculate the cone tip resistance between Df and Df + 15B below the foundation depth
qt = qc + u2 ∙ (1 minus a) = 1250 psi + 0 psi ∙ (1 minus 08) = 1000 psi
Step 2a CPT Material Index
Using Figure 28 the sleeve friction at 15B below the foundation depth is about 16 psi
65
fs 16 psi Fr() = 100 ∙ = 100 ∙ = 13
(qt minus σvo) (1250 psi minus 189 psi)
Iterate to solve for Ic Steps are not shown for brevity
(qt minus σvo)σatm (1250 psi minus 189 psi )145 psi = = Qtn prime (σvoσatm)n (189 psi145 psi)n
prime σvo 189 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Qtn = 703 n = 072 lt 10 Ic = 210
66
Figure 29 Soil type for example problem 3
Use the Ic value to determine the foundation soil formation parameter
Calculate the foundation soil formation parameter The tip resistance is about 1250 psi at 24 feet and
the cone area ratio was determined to be 08 from information provided by the manufacturer
23 23 hs = 28 minus = 28 minus = 078 = SandSilt 15 15 Ic 210
1 + ( ) 1 + ( ) 24 24
67
Step 3 Determine representative values of soil elastic modulus from soil testing and Poissons ratio The
soil elastic modulus was taken as the average value between depths of Df and 15middotB (6 feet and 24 feet)
433 + 557 + 695 + 590 + 501 Es = = 555 tsf = 708 psi
5
ldquoDrained sandsiltrdquo gives a ν = 020
Step 4 Calculate the net cone tip resistance
qtnet = qt minus σvo = 1250 psi minus 189 psi = 12311 psi
Step 5 Calculate the bearing capacity of the sand
In drained sandssilts the bearing capacity is taken as the stress when (sB) = 011 (or 11 foundation width)
minus0345 minus0345 05 s L 600 qmax = hs ∙ qtnet ∙ ( ) ∙ ( ) = 078 ∙ (9795) ∙ (011)05 ∙ ( )
B B 144
= 193 psi = 27860 psf qmax
Assuming a factor of safety (FS) of 3
qmax = 645 psi = 9287 psf
3
Step 6 Calculate settlement
68
119902119898119886119909 0345 2
1 frasl119865119878 L 119904 = 119861 ∙ [ ∙ ∙ ( ) ]
ℎ119904 119902119905119899119890119905 B
2 0345 1 645 psi 600 inches s = 144 inches ∙ [ ∙ ∙ ( ) ] = 18 inches
078 12311 psi 144 inches
Step 7 Determine if q gt qmax
q = 8000 psf lt 9287psf = qmax3
69
CHAPTER 4 DIRECT CPT METHOD FOR DEEP FOUNDATIONS
41 INTRODUCTION
The axial compression capacity (Qtotal) for a single pile includes a side component (Qside) end bearing
component (Qbase) and pile weight (Wpile) as shown in Figure 30 Since piles commonly push through
several layers a summation of the unit side frictions acting on the pile segments must be considered
over the length of the pile While the examples shown in this Guide resemble hand calculations
computer software is more efficient Programming the procedure is possible and represents a practical
method of designing deep foundations However commercial software is frequently available and is
MnDOTrsquos most common method of designing deep foundations
There are upwards of 40 different direct CPT methods that have been developed over the past five
decades to determine a piles axial compression capacity Many of the earliest methods relied on hand-
recorded information where qc data from mechanical CPTs would be collected at 20 cm intervals
whereas the direct CPT method uses scaled penetrometer readings via specified algorithms to obtain
the pile unit side friction and unit end bearing The method that will be used for deep foundation design
is the Modified UniCone method which uses all three readings of the electronic piezocone (qt fs and u2)
while addressing a variety of pile foundation types
Figure 30 Direct CPT evaluation of axial pile capacity
70
The modified UniCone Method is based upon a total of 330 pile load tests (three times the original
Unicone database) that were associated with SCPTu data Originally the UniCone method provided
approximate soil classification in five groups via a chart of qE vs fs (Figure 31) where qE = qt -u2 Later
using the modified approach with a larger data set provided soil sub classifications as shown in Figure
32 This new 9-zone normalized soil behavior type is determined using CPT data in combination with the
CPT Material Index
Figure 31 UniCone Method soil behavior type using CPT (Mayne 2017)
Figure 32 Modified UniCone Method soil behavior type using CPT (Mayne 2017)
71
42 MODIFIED UNICONE METHOD
In order to determine the axial pile capacity using the modified method the first step requires the
determination of geoparameters as shown in Direct CPT Method for Soil Characterization Once the
soil unit weight and CPT Material index are determined for each soil layer then the effective cone
resistance pile unit side friction (fp) and pile end bearing resistance (qb) can be determined
421 Step 1
Work through the steps provided in CPT Method for Soil Characterization until each soil layer is
defined by its CPT material index and soil behavior type using Figure 6 After these steps have been
completed continue to Step 2 to determine qE qb and fp
422 Step 2
Once qt is determined for each layer the effective cone resistance can be calculated using Equation 37
119902119864 = 119902119905 minus 1199062 37
Where (a) qE is the specific value at each elevation along the pile sides for determining fp and (b) at the bottom of the pile qE is averaged in the vicinity of the pile tip from the tip bearing elevation to about one diameter beneath the tip for determining qb
Using CPT material index and qE the pile end bearing resistance is calculated using Equation 38
38 119902119887 = 119902119864 ∙ 10(0325∙119868119888minus1218)
The pile unit side friction is obtained from qE and Ic
39 119891119901 = 119902119864 ∙ 120579119875119879 ∙ 120579119879119862 ∙ 120579119877119860119879119864 ∙ 10(0732∙119868119888minus3605)
Where θPT = coefficient for pile type (084 for bored 102 for jacked 113 for driven piles) θTC = coefficient for loading direction (111 for compression and 085 for tension) θRATE = rate coefficient applied to soils in SBT zone 1 through 7 (109 for constant rate of penetration test and 097 for maintained load tests)
72
423 Step 3
Due to piles extending through multiple layers the unit side components acting on various pile
segments would need to be summed if not using the direct CPT method Since CPT calculates data at
regular intervals of 2 cm to 5 cm along the sides of the pile the average fp in each layer can be used
directly in Equation 40 to obtain the shaft capacity
119876119904119894119889119890 = 119891119901 ∙ 119860119904 40
Where fp = average pile side friction along pile length from eqn 39 As = πmiddotdmiddotH d = pile diameter and H = length embedded below grade
The base capacity for a pile in compression loading is given by Equation 41 For piles in tension (or uplift)
Qbase can be taken as 0
= 119902119887 ∙ 119860119887 41 119876119887119886119904119890 Where
qb = end bearing resistance from eqn 38 Ab = πmiddotd24 (area of a circular pile)
424 Step 4
The final step is to calculate the axial pile capacity using Equation 42
119876119905119900119905119886119897 = 119876119904119894119889119890 + 119876119887119886119904119890 minus 119882119901119894119897119890 42
43 AXIAL PILE DISPLACEMENTS
Movement of pile foundations can be assessed using elastic continuum theory which has been
developed using finite element analyses boundary elements and analytical closed-form solutions In
the case of piles passing through several soil layers the elastic solution can be used by stacking pile
segments (each with its own stiffness) as represented by soil Youngrsquos modulus The use of software is
recommended for pile groups Several available programs such as DEFPIG GROUP and PIGLET can
handle pile groups under axial and lateralmoment loading
73
44 EXAMPLE PROBLEMS
441 Example 5 Direct CPT Methods Axial Pile Capacity
Several piles need to be placed beneath the edge of a building Use the CPT data collected for this site
shown in Figure 34 to determine the axial capacity for one of the piles Assume round steel driven piles
will be used with a diameter of 1275 inches wall thickness of 025 inches and lengths of 80 feet
Concrete will be used to fill the piles To solve for all geoparameters use the Direct CPT Method for Soil
Characterization section All sand layers can be assumed ldquodrainedrdquo with ν = 02
Figure 33 Deep foundation end bearing pile diagram
74
Figure 34 CPT data from Minnesota for example problem 5
75
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 35 Soil layers using ldquorules of thumbrdquo for pile capacity example using direct CPT method
Layer 1
From Figure 35 fs = 18 psi taken as a representative value of the layer
76
18 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1210 pcf
145 psi
Layer 2
From Figure 35 fs = 12 psi taken as a representative value of the layer
12psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 3
From Figure 35 fs = 20 psi taken as a representative value of the layer
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
CPT Material Index
Layer 1
From Figure 35 fs = 18 psi and qc = 3000 psi
qt = qc + u2 ∙ (1 minus a) = 3000 psi + 20 psi ∙ (1 minus 08) = 3004 psi
∙ 4 feet = 1219 pcf ∙ 4 feet = 4877 psf = 34 psi σvo = γt
fs 18 psi Fr() = 100 ∙ = 100 ∙ = 060
(qt minus σvo) (3004 psi minus 34 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
77
uo = 0 psi
prime σvo = σvo minus uo = 34 psi minus 0 psi = 34 psi
(qt minus σvo)σatm (3004 psi minus 34 psi )145 psi = = Qtn prime )n (σvoσatm (34 psi145 psi)n
prime σvo 34 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(060)]2
Qtn = 3593 n = 038 lt 10 Ic = 14 (ie sand)
Layer 2
From Figure 35 fs = 12 psi and qc = 500 psi
qt = qc + u2 ∙ (1 minus a) = 500 psi + 40 psi ∙ (1 minus 08) = 508 psi
= 4838 psf + γt ∙ 45 feet = 4838 psf + 1172 pcf ∙ 45 feet = 5756 psf = 400 psi σvo
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 260
(qt minus σvo) (508 psi minus 400 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
78
uo = 0 psi
prime σvo = σvo minus uo = 400 psi minus 0 psi = 400 psi
(qt minus σvo)σatm (508 psi minus 400 psi )145 psi = = Qtn prime (σvoσatm)n (400 psi145 psi)n
prime σvo 400 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(260)]2
Qtn = 117 n = 10 le 10 Ic = 29 (ie clayey silt)
Layer 3
From Figure 35 fs = 20 psi and qc = 5000 psi
qt = qc + u2 ∙ (1 minus a) = 5000 psi + 0 psi ∙ (1 minus 08) = 5000 psi
= 5756 psf + γt ∙ 6 feet = 5756 psf + 1219 pcf ∙ 6 feet = 6488 psf = 451 psi σvo
fs 20 psi Fr() = 100 ∙ = 100 ∙ = 040
(qt minus σvo) (5000 psi minus 451 psi)
79
Step 2 Iterate to solve for Ic Steps are not shown for brevity
prime σvo = σvo minus uo = 451 psi minus 0 psi = 451 psi
(qt minus σvo)σatm (5000 psi minus 451 psi )145 psi = = Qtn prime )n (σvoσatm (451 psi145 psi)n
prime σvo 451 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(040)]2
= 1801 n = 06 lt 10 = 15 (ie sand) Qtn Ic
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic (14) Qtn (359) and Fr (06) the first layer is defined as a ldquoGravelly Sandrdquo
from Figure 36
80
Figure 36 Soil layer 1 using SBT method
81
Layer 2
Based on values of Ic (29) Qtn (117) and Fr (26) the second layer is defined as a ldquoSilty Mixrdquo
from Figure 37
Figure 37 Soil layer 2 using SBT method
82
Layer 3
Based on values of Ic (15) Qtn (180) and Fr (04) the third layer is defined as a ldquoSandrdquo from
Figure 38
Figure 38 Soil layer 3 using SBT method
Effective Cone Resistance
Layer 1
83
qE = qt minus u2 = 3004 psi minus 20 psi = 2984 psi
Layer 2
qE = qt minus u2 = 508 psi minus 40 psi = 468 psi
Layer 3
qE = qt minus u2 = 5000 psi minus 0 psi = 5000 psi
Pile End Bearing Resistance
Layer 3
= qE ∙ 10(0325∙Icminus1218) qb = 5000 psi ∙ 10(0325∙15minus1218) = 9085 psi
Pile Unit Side Friction
Layer 1
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 2984 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙14minus3605) = 99 psi
Layer 2
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 468 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙29minus3605) = 211 psi
Layer 3
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 5000 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙15minus3605) = 202 psi
84
Axial Pile Capacity
The side capacity is based on pile design as shown in Figure 39
Layer 1
Qside = fp ∙ As = 99 psi ∙ (π ∙ d ∙ L) = 99 psi ∙ (π ∙ 1275 in ∙ 48 in) = 19064 lb
Layer 2
Qside = fp ∙ As = 211 psi ∙ (π ∙ d ∙ L) = 211 psi ∙ (π ∙ 1275 in ∙ 588 in) = 457122 lb
Layer 3
Qside = fp ∙ As = 202 psi ∙ (π ∙ d ∙ L) = 202 psi ∙ (π ∙ 1275 in ∙ 240 in) = 58170 lb
Figure 39 Soil layering compared to pilings
85
End Bearing Resistance in Layer 3
d2 1275 in2
Qbase = qb ∙ Ab = 9085 psi ∙ (π ∙ ) = 9085 psi ∙ (π ∙ ) = 115932 lb 4 4
Wp = (γsteel ∙ Asteel ∙ Depth) + (γsteel ∙ Aconc ∙ Depth)
Wp = (490 pcf ∙ 003 ft2 ∙ 55 ft) + (150 pcf ∙ 085 ft2 ∙ 55 ft) = 7724 lbs
Layer 3
Qtotal = ΣQside + Qblayer3 minus Wp
Qtotal = (19064 + 45713 + 58170) lbs + 115932 lbs minus 7724 lbs = 642563 lbs
= 642 kips Qtotal
Table 8 Summary of selected and calculated parameters
Layer L (in) qc
(psi)
fs
(psi)
Fr Ic Qtn qE
(psi)
qb
(psi)
Qbase
(lb)
fp
(psi)
Qside
(lb)
Qtotal
(kip)
1 48 3000 18 06 14 356 2984 NA NA 99 19064
2 588 500 12 26 29 117 468 NA NA 211 457122
3 240 5000 20 04 15 180 5000 9085 115932 202 58170
Total 115932 534355 642
86
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Agaiby SA (2018) Advancements in the interpretation of seismic piezocone tests in clays and other geomaterials PhD dissertation School of Civil amp Environmental Engineering Georgia Institute of Technology Atlanta GA
Agaiby SS and Mayne PW (2018) Interpretation of piezocone penetration and dissipation tests in sensitive Leda Clay at Gloucester Test Site Canadian Geotechnical Journal (accepted for publication 08 March 2018) CGJ-2017-0388
Amar S Baguelin F Canepa Y and Frank R (1998) New design rules for the bearing capacity of shallow foundations based on Menard PMT Geotechnical Site Characterization Vol 2 (Proc ISC-1 Atlanta) 727-733
American Petroleum Institute (API) (1981) Planning designing and constructing fixed offshore platforms Recommended practice API-RP2A 17th edition Washington DC API
American Petroleum Institute (API) (1987) Planning designing and constructing fixed offshore platforms Recommended practice API-RP2A 18th edition Washington DC API
American Petroleum Institute (API) (1989) Planning designing and constructing fixed offshore platforms Recommended practice API-RP2A 19th edition Washington DC API
Andersen KH and Stenhamar P (1982) Static plate loading tests on overconsolidated clay Journal of the Geotechnical Engineering Division ASCE 108 (GT7) 919-934
Basu D and Salgado R (2012) Load and resistance factor design of drilled shafts in sands Journal of Geotechnical amp Geoenvironmental Engineering 138 (12) 1455-1469
Berardi R Jamiolkowski M and Lancellotta R (1991) Settlement of shallow foundations in sands selection of stiffness on the basis of penetration resistance Geotechnical Engineering Congress Vol 1 185-200 (Proc GSP-27 Boulder) Reston Virginia ASCE
Brand EW Muktabhant C and Taechathummarak A (1972) Load tests on small foundations in soft clay Performance of Earth and Earth-Supported Structures Vol 1 Part 2 903-928 (Proc PC) ASCE Reston Virginia ASCE
Briaud JL and Gibbens RM (1999) Behavior of five large spread footings in sand Journal of Geotechnical amp Geoenvironmental Engineering 125 (9) 787-797
Briaud JL (2007) Spread footings in sand load-settlement curve approach Journal of Geotechnical amp Geoenvironmental Engineering 133 (8) 905-920
Brown DA Turner JP and Castelli RJ (2010) Drilled Shafts Construction Procedures and LRFD Design Methods Geotechnical Engineering Circular GEC 10 Report No FHWA NHI-10-016 Washington DC National Highway Institute US DOT
87
Burland J B (1973) Shaft Friction of Piles in Clay -A Simple Fundamental Approach Ground Eng 6(3) 30-42
Cai G Liu S and Puppala A (2012) Reliability assessment of CPTu-based pile capacity predictions in soft clay deposits Engineering Geology 84-91
Canadian Geotechnical Society (1992) Canadian Foundation Engineering Manual Third Edition Vancouver British Columbia BiTech Publishers
Cerato AB and Lutenegger AJ (2007) Scale effects of shallow foundation bearing capacity on granular material Journal of Geotechnical amp Geoenvironmental Engineering 133 (10) 1192-1201
Chen BS-Y and Mayne PW (1996) Statistical relationships between piezocone measurements and stress history of clays Canadian Geotechnical Journal 33 (3) 488-498
Cho W and Finno RJ (2010_ Stress-strain responses of block samples of compressible Chicago glacial clays Journal of Geotechnical amp Geoenvironmental Engineering 136 (1) 178-188
Chung SF Randolph MF and Schneider JA (2006) Effect of penetration rate on penetrometer resistance in clay Journal of Geotechnical amp Geoenvironmental Engineering 132 (9) 1188-1196
Clausen CJF Aas PM and Karlsrud K (2005) Bearing capacity of driven piles in sand the NGI approach Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis Group
Consoli NC Schnaid F and Milititsky J (1998) Interpretation of plate load tests on residual soil site Journal of Geotecnical amp Geoenvironmental Engingeering 124 (9) 857-867
Dasenbrock DD (2006) Assessment of pile capacity by static and dynamic methods to reconcile design predictions with observed performance Proc 54th Minnesota Geotechnical Engineering Conference 95-108 edited by Labuz J Univ of Minnesota St Paul
DeBeer E and Martens A (1957) Method of computation of an upper limit for the influence of heterogeneity of sand layers on the settlement of bridges 275-282 Proc ICSMFE-4 Vol 1 London ICSMFE
DeBeer EE (1965) Bearing capacity and settlement of shallow foundations on sand Proc Symposium on Bearing Capacity and Settlement of Foundations 15-33 Duke University Durham NC
Decourt L (1999) Behavior of foundations under working load conditions Proc PCSMGE-11 Foz do Iguassu Brazil Vol 4 453-487
Demers D and Leroueil S (2002) Evaluation of preconsolidation pressure and the overconsolidation ratio from piezocone tests of clay deposits in Quebec Canadian Geotechnical Journal 39 (1) 174-192
Eslaamizaad S and Robertson PK (1996) Cone penetration test to evaluate bearing capacity of foundations in sand 429-438 Proc CGCFG-49 Vol 1 St Johns Newfoundland
Eslami A Alimirzaei M Aflaki E and Molaabasi H (2017) Deltaic soil behavior classification using CPTu records Marine Georesources amp Geotechnology 35 (1) 62-79
88
Eslami A and Gholami M (2005) Bearing capacity analysis of shallow foundations from CPT data 1463-1466 Proc ICSMGE- 16 Vol 3 (Osaka) Millpress Rotterdam
Eslami A and Gholami M (2006) Analytical model for the ultimate bearing capacity of foundations from cone resistance Scientia Iranica 13 (3) 223-233
Eslami A and Fellenius BH (1997) Pile capacity by direct CPT and CPTu methods applied to 102 case histories Canadian Geotechnical Journal 34 (6) 886-904
Fellenius BH (2009) Basics of Foundation Design Vancouver BiTech Publishers
Fellenius BH and Altaee A (1994) Stress and settlement of footings in sand Vertical and Horizontal Deformations of Foundations and Embankments 2 1760-1773
Fellenius BH and Eslami A (2000) Soil profile interpreted from CPTu data Proc Geotechnical Engineering Conference Year 2000 Geotechnics Asian Inst of Technology Bangkok Thailand Available from wwwfelleniusnet
Finno R J and Chung C K (1992) Stress-strain-strength responses of compressible Chicago glacial clays Journal of Geotechnical Engineering 118 (10) 1607ndash1625
Fleming K Weltman A Randolph M and Elson K (2009) Piling Engineering Third Edition London Taylor amp Francis Group
Frank R and Magnan J-P (1995) Cone penetration testing in France national report Proceedings International Symposium on Cone Penetration Testing 3 (CPT95 Linkoumlping) Swedish Geotechnical Society Report 3(95) 147-156
Gavin K Adekunte A and OKelly B (2009) A field investigation of vertical footing response on sand Geotechnical Engineering 162 257-267
Hegazy YA (1998) Delineating geostratigraphy by cluster analysis of piezocone data PhD dissertation School of Civil amp Environmental Engineering Georgia Institute of Technology Atlanta GA
Holtz RD Kovacs WD and Sheehan TC (2011) An Introduction to Geotechnical Engineering 2nd Edition London Pearson Publishing
Hunt CE Pestana JM Bray JD and Riemer M (2002) Effect of pile driving on static and dynamic properties of soft clay Journal of Geotechnical amp Geoenvironmental Engineering 128 (1) 13-24
Jamiolkowski M Ladd CC Germaine JT and Lancellotta R (1985) New developments in field and lab testing of soils Proc ICSMFE-11 1 57-154
Jardine R Chow F Overy R and Standing J (2005) ICP design methods for driven piles in sands and clays London Thomas Telford Publishing
Jardine RJ Lehane BM Smith P and Gildea P (1995) Vertical loading experiments on rigid pad foundations at Bothkennar Geotechnique 45 (4) 573-597
89
Karlsrud K (2012) Prediction of load-displacement behavior and capacity of axially-loaded piles in clay based on analyses and interpretation of pile load test results PhD Dissertation Norwegian Univ Science amp Technology Trondheim Norway
Karlsrud K Clausen CJF and Aas PM (2005) Bearing capacity of driven piles in clay the NGI approach Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis
Karlsrud K Lunne T Kort DA and Strandvik S (2001) CPTU correlations for clays Proc 16th ICSMGE 2 683-702
Kimura T Kusakabe O and Saitoh K (1985) Geotechnical model tests of bearing capacity problems in a centrifuge Geotechnique 35 (1) 33-45
Knappett JA and Craig RF (2012) Craigs Soil Mechanics 8th Edition London Spon Press Taylor amp Francis
Kolk HJ and Velde EVD (1996) A reliable method to determine friction capacity of piles driven into clays Paper No 7993 Proc Offshore Technology Conference Houston
Kolk HJ Baaijens AE and Senders M (2005) Design criteria for pipe piles in silica sands Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis Group
Kulhawy FH (2004) On the axial behavior of drilled shafts Geosupport 2004 1-18
Kulhawy FH and Jackson CS (1989) Foundation Engineering Current Principles amp Practices 2 (GSP 22) Reston VA ASCE
Kulhawy FH and Mayne PW (1990) Manual for Estimating Soil Properties for Foundation Design EPRI Report EL-6800 Electric Power Research Institute Palo Alto 306 page Download from wwwepricom
Kulhawy FH Trautmann CH Beech JF ORourke TD and McGuire W (1983) Transmission Line Structure Foundations for Uplift-Compression Loading Report EL-2870 Palo Alto CA Electric Power Research Institute wwwepricom
Ku T and Mayne PW (2015) In-situ lateral stress coefficient (K0) from shear wave velocity measurements in soils Journal of Geotechnical amp Geoenvironmental Engineering 141 (12) 101061(ASCE) GT1943-56060001354
Ladd CC and DeGroot DJ (2003) Recommended practice for soft ground site characterization Soil amp Rock America 2003 3-57
Larsson R (1997) Investigations and Load Tests in Silty Soils SGI Report R-54 Linkoumlping Swedish Geotechnical Institute
Larsson R (2001) Investigations and Load Tests in Clay Till SGI Report R-59 Linkoumlping Swedish Geotechnical Institute
Lee J and Salgado R (2005) Estimation of bearing capacity of circular footings on sands based on cone penetration tests Journal of Geotechnical amp Geoenvironmental Engineering 131 (4) 442-452
90
Lehane BM (2003) Vertically loaded shallow foundation on soft clayey silt Geotechnical Engineering 156 (1) Thomas Telford London 17-26
Lehane BM (2013) Relating foundation capacity in sands to CPT qc Geotechnical amp Geophysical Site Characterization 1 London Taylor amp Francis Group
Lehane BM Doherty JP and Schneider JA (2008) Settlement prediction for footings on sand Deformational Characteristics of Geomaterials (IS-Atlanta) 1 Rotterdam Millpress
Lehane BM Li Y and Williams R (2013) Shaft capacity of displacement piles in clay using the CPT Journal of Geotechnical amp Geoenvironmental Engineering 139 (2) 253-266
Lehane BM Schneider JA and Xu X (2005) The UWA-05 method for prediction of axial capacity of driven piles in sand Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis
Low HE Lunne T Andersen KH Sjursen MA Li X and Randolph MF 2010 Estimation of intact and remoulded undrained shear strengths from penetration tests in soft clays Geotechnique 60 (11) 843-859
Lunne T Robertson PK and Powell JJM (1997) Cone Penetration Testing in Geotechnical Practice London RoutledgeBlackie Academic
Lutenegger AJ and DeGroot DJ (1995) Settlement of Shallow Foundations on Granular Soils Report Contract 6332 Task Order 4 submitted to Massachusetts Highway Dept by University of Mass-Amherst Amherst MA
Lutenegger AJ and Adams MT (2003) Characteristic load-settlement behavior of shallow foundations Proc FONDSUP 2 381-392
Mase T and Hashiguchi K (2009) Numerical analysis of footing settlement problem by subloading surface model Soils amp Foundations 49 (2) 207-220
Mayne PW (1987) Determining preconsolidation stress and penetration pore pressures from DMT contact pressures Geotechnical Testing Journal 10(3) 146-150
Mayne PW (1991) Determination of OCR in Clays by Piezocone Tests Using Cavity Expansion and Critical State Concepts Soils and Foundations 31 (2) 65-76
Mayne PW (2007) NCHRP Synthesis 368 on Cone Penetration Test Washington DC Transportation Research Board National Academies Press wwwtrborg
Mayne PW (2009) Geoengineering Design Using the Cone Penetration Test Richmond BC Canada ConeTec Inc
Mayne PW (2014) Keynote Interpretation of geotechnical parameters from seismic piezocone tests Proceedings 3rd Intl Symp on Cone Penetration Testing 47-73
Mayne PW (2016) Evaluating effective stress parameters and undrained shear strengths of soft-firm clays from CPT and DMT Australian Geomechanics Journal 51 (4) 27-55
91
Mayne PW (2017) Stress history of soils from cone penetration tests 34th Manual Rocha Lecture Soils amp Rocks Vol 40 (3) Saouml Paulo ABMS
Mayne PW Christopher B Berg R and DeJong J (2002) Subsurface Investigations -Geotechnical Site Characterization Publication No FHWA-NHI-01-031 Washington DC National Highway Institute Federal Highway Administration
Mayne PW Coop MR Springman S Huang A-B and Zornberg J (2009) Geomaterial Behavior and Testing Proc 17th Intl Conf Soil Mechanics amp Geotechnical Engineering 4 Rotterdam MillpressIOS Press
Mayne PW and Illingworth F (2010) Direct CPT method for footings on sand using a database approach Proc ISCPT-2 3 315-322
Mayne PW and Niazi F (2017) Recent developments and applications in field investigations for deep foundations Proceedings 3rd Bolivian Conference on Deep Foundations 1 141-160
Mayne PW Peuchen J and Baltoukas D (2015) Piezocone evaluation of undrained strength in soft to firm offshore clays Frontiers in Offshore Geotechnics III 2 1091-1096
Mayne PW and Poulos HG (1999) Approximate displacement influence factors for elastic shallow foundation systems ASCE Journal of Geotechnical amp Geoenvironmental Engineering 125 (6) 453-460
Mayne PW Uzielli M and Illingworth F (2012) Shallow footing response on sands using a direct method based on cone penetration tests Full Scale Testing and Foundation Design (GSP 227) Reston Virginia ASCE
Mayne PW and Woeller DJ (2014) Direct CPT method for footings on sands silts and clays Geocharacterization for Modeling and Sustainability (GSP 234) 1983-1997
McClelland B (1974) Design of Deep Penetration Piles for Ocean Structures Journal Geot Eng Div 100(GT7) 705-747
Mesri G (1975) Discussion on new design procedure for stability of soft clays ASCE Journal of the Geotechnical Engineering Division 101(GT4) pp 409-412
Meyerhof GG (1956) Penetration tests and bearing capacity of cohesionless soils Journal of Soil Mechanics amp Foundations Division 82 (SM1) 1-19
Meyerhof GG (1974) General Report Penetration testing outside Europe Proceedings ESOPT-1 Vol 21 Swedish Geotechnical Society Stockholm
Meyerhof GG (1976) Bearing capacity and settlement of pile foundations Journal of Geotechnical Engineering Division 102(3) 197-228
Missiaen T Verhegge J Heirman K and Crobe P (2015) Potential of cone penetrating testing for mapping deeply buried palaeolandscapes in the context of archaeological surveys in polder areas Journal of Archaeological Science 55 174-187
92
Mlynarek Z Wierzbicki J Goglik S and Bogucki M (2014) Shear strength and deformation parameters of peat and gyttja from CPTu DMT and VST Proceedings 5th International Workshop on CPTu and DMT in soft clays and organic soils 193-210 Poznan Poland Polish Committee on Geotechnics Menard Polska Tensar Internationa
Nejaim PF Jannuzzi GMF and Danziger FAB (2016) Soil behavior type of the Sarapuiacute II test site Geotechnical amp Geophysical Site Characterization 5 1009-1014
Niazi FS and Mayne PW (2013) Cone penetration test based direct methods for evaluating static axial capacity of single piles Geotechnical and Geological Engineering 31 (4) 979-1009
Niazi FS and Mayne PW (2015) Enhanced Unicone expressions for axial pile capacity evaluation from piezocone tests Proc International Foundations Conference amp Equipment Expo RestonVA ASCE
Niazi FS and Mayne PW (2016) CPTu-based enhanced UniCone method for pile capacity Engineering Geology 212 21-34
Nowacki F Karlsrud K and Sparrevik P (1996) Comparison of recent tests on OC clay and implications for design Large Scale Pile Tests in Clay London Thomas Telford
ONeill MW (2001) Side resistance in piles and drilled shafts Journal of Geotechnical amp Geoenvironmental Engineering 127 (1) 3-16
Ouyang Z and Mayne PW (2018) Effective friction angle of clays and silts from piezocone Canadian Geotechnical Journal in press doiorg101139cgj-2017-0451
Paikowsky SG Canniff MC Lesny K Kisse A Amatya S and Muganga R (2010) LRFD Design and Construction of Shallow Foundations for Highway Bridge Structures NCHRP Report 651 Washington DC National Cooperative Highway Research Program Transportation Research Board
Pestana JM Hunt CE and Bray JD (2002) Soil deformation and excess pore pressure filed around a closed-ended pile Journal of Geotechnical amp Geoenvironmental Engineering 128 (1) 1-12
Poulos HG and Davis EH (1974) Elastic Solutions for Soil and Rock Mechanics New York Wiley amp Sons
Poulos HG and Davis E (1980) Pile Foundation Analysis amp Design New York John Wiley amp Sons
Powell JJM and Lunne T (2005) Use of CPTU data in clays and fine-grained soils Studia Geotechnica et Mechanica XXVII(3) 29-66
Randolph MF (2003) Rankine Lecture Science and empiricism in pile foundation design Geotechnique 53 (10) 847-875
Robertson PK (1990 1991) Soil classification using the cone penetration test Canadian Geotechnical Journal 27 (1) 151-158 Closure 28 (1) 176-178
Robertson PK (1991) Estimation of foundation settlements in sand from CPT Geotechnical Engineering Congress 2 Reston Virginia ASCE
93
Robertson PK (2009) Cone penetration testing a unified approach Canadian Geotechnical J 46 (11) 1337-1355
Robertson PK (2009a) Performance based earthquake design using the CPT Keynote Lecture International Conference on Performance-Based Design in Earthquake Geotechnical Engineering IS Tokyo Tsukuba Japan
Robertson PK (2009b) Interpretation of cone penetration tests a unified approach Canadian Geotechnical Journal 46 (11) 1337-1355
Robertson PK and Cabal KL (2007) Guide to cone penetration testing Second Edition Signal Hill California Gregg In-Situ
Robertson PK and Cabal KL (2015) Guide to Cone Penetration Testing for Geotechnical Engineering 6th Edition Signal Hill California Gregg Drilling amp Testing Inc
Schmertmann JH (1970) Static cone to compute static settlement over sand Journal of the Soil Mechanics and Foundations Division 95 (SM3) 1011-1043
Schmertmann JH (1978) Guidelines for cone penetration test performance and design Report FHWA-TS-78-209 Washington DC Federal Highway Administration
Schnaid F Wood WR Smith AKC and Jubb P (1993) Investigation of bearing capacity and settlements of soft clay deposits at Shellhaven Predictive Soil Mechanics London Thomas Telford
Schneider JA Xu X and Lehane BM (2008) Database assessment of CPT-based design methods for axial capacity of driven piles in siliceous sands J Geotechnical and Geoenvironmental Engineering 134 (9) 1227-1244
Semple D (1980) Recent Developments in the Design amp Construction of Piles London Institution of Civil Engineers
Senneset K Sandven R amp Janbu N (1989) Evaluation of soil parameters from piezocone tests In-Situ Testing of Soil Properties for Transportation Transportation Research Record 1235 24-37
Shahri AA Melehmir A and Juhlin C (2015) Soil classification analysis based on piezocone penetration test data Engineering Geology 189 32-47
Stuedlein AW and Holtz RD (2010) Undrained displacement behavior of spread footings in clay The Art of Foundation Engineering Practice GSP 198 Reston Virginia ASCE
Takesue K Sasao H and Matsumoto T (1998) Correlation between ultimate pile skin friction and CPT data Geotechnical Site Characterization 2 1177ndash1182
Tand KE Funegard EG and Briaud J-L (1986) Bearing capacity of footings on clay CPT method Use of In-Situ Tests in Geotechnical Engineering GSP 6 1017-1033
Tand KE Warden PE and Funegaringrd EG (1995) Predicted-measured bearing capacity of shallow footings on sand PISCPT 2 589-594
Thomas D (1968) Deep soundings test results and the settlement of spread footings on normally consolidated sands Geotechnique 18 (2) 472-488
94
Tomlinson MJ (1957) The adhesion of piles driven into clay soils Proc 4th Intl Conf on Soil Mechanics and Foundation Engineering 2 66-71
Tomlinson MJ (1986) Foundation Design amp Construction London Longman Scientific
Tomlinson MJ (1987) Pile Design and Construction Practice London Viewpoint Publication
Trofimenkov JG (1974) Penetration testing in the USSR Proceedings ESOPT-1 1 147-154
Tumay MT Hatipkarasulu Y Marx ER and Cotton B (2013) CPTPCPT-based organic material profiling Proc 18th Intl Conf on Soil Mechanics amp Geotechnical Engineering Paris 633-636
Uzielli M and Mayne PW (2011) Statistical characterization and stochastic simulation of load-displacement behavior of shallow footings Proc Georisk 2011 Geotechnical Risk Management amp Assessment (GSP 224) 672-679
Uzielli M and Mayne PW (2012) Load-displacement uncertainty of vertically-loaded shallow footings on sands and effects on probabilistic settlement estimation Georisk Assessment and Management of Risk for Engineered Systems and GeoHazards 6 (1) 50-69
Valsson SM (2016) Detecting quick clay with CPTu Proceedings 17th Nordic Geotechnical Meeting Reykajavik Iceland Challenges in Nordic Geotechnics
Van Dijk BFJ and Kolk HJ (2011) CPT-based design method for axial capacity of offshore piles in clay Frontiers in Offshore Geotechnics II 555-560
Vesić AS (1975) Bearing capacity of shallow foundations Foundation Engineering Handbook New York Van Nostrand Reinhold Publishers
Vesic AS (1977) NCHRP Synthesis of Highway Practice 42 Design of Pile Foundations Washington DC Transportation Research Board
Yu F and Yang J (2012) Base capacity of open-ended steel pipe piles in sand J Geotechnical and Geoenvironmental Engineering 138 (9) 1116-1128
Zawrzykraj P Rydelek P and Bakowska A (2017) Geoengineering properties of Eemian peats from Radsymin (central Poland) in the light of static cone penetration and dilatometer tests Engineering Geology 226 290-300
95
APPENDIX A
List of Figures
Figure A1 Conventional methods versus direct CPT approach to shallow foundation response A-3
Figure A2 Direct bearing capacity relationship for embedded square and strip footings situated on clean
sands (after Schmertmann 1978) A-6
Figure A3 Direct CPT bearing factor Rk as function of footing width B and embedment depth from finite
element analyses (Tand et al 1995) A-6
Figure A4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio and sand
consistency (Eslaamizaad amp Robertson 1996) A-7
Figure A5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp Salgado (2005) in
terms of footing size relative density and base settlement A-8
Figure A6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and normalized
cone tip resistance (after Eslami amp Gholami 2005) A-8
Figure A7 Direct relationship between ultimate bearing stress in clays and measured cone tip resistance
(Schmertmann 1978) A-10
Figure A8 Direct relationship between ultimate foundation bearing stress in clays and net cone tip
resistance (after Tand et al 1986) A-11
Figure A9 Measured load-displacements for each of the five large footings at TAMU (Briaud amp Gibbens
1994) sand site A-12
Figure A10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU (Briaud amp
Gibbens 1994) footings A-13
Figure A11 Characteristic stress versus square root normalized displacements for TAMU footings A-15
Figure A12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sand A-16
Figure A13 Summary of sand formation factors with corresponding CPT resistances A-19
Figure A14 Normalized foundation stresses vs square root of normalized displacement for spread
footings on sand (modified after Mayne et al 2012) A-20
Figure A15 Definitions of capacity from three types of observed foundation load test responses
(modified after Kulhawy 2004) A-21
Figure A16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footings A-22
A-1
Figure A17 Measured load vs displacement curves for 3 shallow foundations on clay at Baytown Texas
(Stuedlein amp Holtz 2010) A-25
Figure A18 Characteristic stress vs square root of normalized displacement for all three foundations at
Baytown site (Stuedlein amp Holtz 2010) A-26
Figure A19 Normalized foundation stress vs square root of normalized displacements A-27
Figure A20 Applied foundation stress vs CPT-calculated stress for 67 footings A-28
Figure A21 Applied footing stress vs CPT calculated stress on logarithmic scales A-29
Figure A22 Summary graph and equations of direct CPT method for evaluating the vertical A-30
Figure A23 Foundation soil formation parameter hs versus CPT material index Ic A-32
Figure A24 Bearing capacity ratio qmaxqnet versus CPT material index Ic A-33
Figure A25 Influence factor for rectangular foundations from elastic theory solution (Mayne amp
Dasenbrock 2017) A-34
List of Tables
Table A1 Direct CPT methods for bearing capacity of footings on clean sands A-4
Table A2 Direct CPT methods for bearing capacity of footings on clays A-9
Table A3 Summary of large footings soil conditions and reference sources of database A-17
Table A4 Sources for shallow foundation performace A-34
A-2
A1 INTERPRETATION OF SOIL PARAMETERS FROM CONE
PENETRATION TESTS
A11 Introduction
Geotechnical site characterization is an important first step towards the evaluation of
subsurface conditions and determination of soil layering geomaterial classification and the
evaluation of soil engineering parameters for the analysis and design of foundations retaining
walls tunnels excavations embankments and slope stability Increasing use of the electronic
cone penetrometer in highway applications is prominent in the USA because the results are
obtained much faster and less expensive than traditional methods that rely on rotary drilling
augering sampling and laboratory testing Moreover the cone penetration test (CPT) provides
at least three independent and continuous readings on soil behavior that are digitally recorded
and fully available at the end of the sounding As such the reliable and consistent
interpretation of CPT data is important since civil engineering works will best realize the
efficiency and economy of this technology in constructed facilities for county state and
interstate projects
A111 Cone Penetration Testing
The cone penetrometer is an electronically-instrumented steel probe that is vertically pushed into the
ground at constant rate of 20 mms (08 insec) using a hydraulic system The penetrometer is outfitted
with load cells pressure transducers and inclinometers that measure at least three readings
continuously with depth (a) cone tip resistance qt (b) sleeve friction fs and (c) porewater pressure u2
With the latter reading the device is called a piezocone thus the designation CPTu is used to indicate
porewater pressure Additional sensors are available that can be used to measure inclination
resistivity pH temperature shear wave velocity and other readings if desired
Figure A1 shows a schematic rendering of the cone penetration test that is performed in the
field per ASTM D 5778 procedures The hydraulic system is nominally rated at 180 kN (20 tons)
capacity and often mounted on a truck but can also be positioned on tracked vehicles or
anchored frames Figure A2 shows a MnDOT cone truck that uses the full dead weight of the
rig for the hydraulic reaction forces which are applied at the center of the vehicle The cab is
enclosed so that soundings may proceed during inclement weather and the data acquisition
system and operator are protected
Figure A 1 Setup and procedures for co ne penetration testing (CPT)
A-4
Figure A2 CPT truck-mounted rig used by MnDOT
A representative sounding taken at the I-35 test site just northeast of Saint Paul is presented in Figure
A3 Here the separate profiles of qt fs and u2 with depth are shown in side-by-side plots In the last
graph the hydrostatic porewater pressure (uo) due to the groundwater table is indicated by the blue
dashed line
These readings are used to interpret the soil layers types and properties of the ground as
discussed in the following sections
A-5
Figure A3 Representative CPT sounding at I-35 test site northeast of St Paul MN showing profiles of
(a) cone resistance (b) sleeve friction and (c) penetration porewater pressures
A112 Soil Parameter Interpretation
A variety of soil engineering parameters can be interpreted from CPT results on the basis of
theoretical frameworks analytical models or numerical simulations otherwise by empirical
methods based on correlations and statistics of databases Figure A4 shows a conceptual
utilization of the readings from the CPT for interpretation of geostratigraphy compressibility
flow characteristics and stress-strain-strength behavior of soils
Table A1 lists a selection of geoparameters that have been addressed for these purposes The
most common or useful relationships will be discussed in subsequent sections and reference is
made to other sources (Lunne et al 1997 Mayne 2007 Robertson amp Cabal 2015) for additional
details
A-6
Figure A4 Conceptual post-processing of CPT results for geoparameter evaluation
Table A1 List of common geoparameters interpreted from CPT data
Symbol Parameter Remarks Notes
SBT Soil behavior type (SBT)
Ic CPT material index
γt Unit weight
ρt Mass Density
σvo Overburden stress
u0 Hydrostatic pressure
A-7
σvo Effective stress
Table A1 Continued
Symbol Parameter Remarks Notes
σp Preconsolidation stress σp = 033 (qnet)m where m = fctn (Ic)
(or Effective Yield Stress) where stresses are in kPa
YSR Yield stress ratio YSR = σpσvo (formerly OCR)
c Effective cohesion intercept
ϕ Effective friction angle
su Undrained shear strength
D Constrained modulus
G0 Small-strain modulus
G Shear modulus
E Youngs modulus
cvh Coefficient of consolidation
K Bulk modulus
A-8
LI
K0 Lateral stress coefficient
k Hydraulic conductivity
Liquefaction index
A12 Geostratigraphy and Soil Behavioral Type (SBT)
In routine CPT soil samples are not normally taken and thus the measured stress friction andor
porewater pressure readings are used to infer the types of soils This can be accomplished using
three basic approaches
a Rules of Thumb or approximate guidelines for a quick visual assessment
b Soil Behavioral Type (SBT) Charts that are based on either the three readings (qt fs u2) net readings
including net cone resistance (qnet = (qt-σvo) effective cone resistance (qE = qt - u2) friction ratio (Rf =
100middotfsqt) and excess porewater pressures or using normalized CPT readings such as Q F Bq or U
c Probabilistic methods where the CPT readings have been calibrated from lab tests on
recovered soil samples (eg Tumay et al 2013)
A121 Approximate Rules of Thumb
The approximate rules of thumb provide simple guidelines for soil type and suggest that sands are
identified when qt gt 5 MPa and u2 asymp u0 whereas intact clays are prevalent when qt lt 5 MPa and u2 gt u0
(Mayne et al 2002) The magnitude of porewater pressures reflects the consistency of the intact clay
such that soft (u2 asymp 2middotu0) firm (u2 asymp 4middotu0) stiff (u2 asymp 8middotu0) and hard (u2 asymp 20middotu0) However for fissured
overconsolidated clays the measured porewater pressures are less than hydrostatic in fact often
negative u2 lt 0 On land negative porewater pressure readings at the shoulder filter element of the
piezocone should be greater than -1 bar ie u2 gt -100 kPa (-147 psi) Also for clean sands the friction
ratio (FR = 100middotfsqt lt 1) while for insensitive clays (FR gt 4)
A-9
Figure A5 presents a 36-m deep piezocone sounding from the MnDOT Wakota Bridge site which crosses
the Mississippi River at I-494 (Dasenbrock 2006) The qt and u2 readings have been annotated using the
aforementioned rules of thumb indicating the predominance of sands at this site As noted there are
five interbedded clay layers evident at depths of 1 m 3-4 m 7 - 12 m 17 m and 22-27 m
Figure A5 Interpretation of soil types from CPT data taken from the Wakota Bridge on I-494 using
approximate rules of thumb
A122 Soil Behavioral Type Charts
The most common approach for identification of soil types is based on soil behavioral charts
Reviews of several chart methods are provided elsewhere (Kulhawy amp Mayne 1990 Fellenius amp
Eslami 2000 Shahri et al 2015) Current popularity favors the 9-zone classification system to
evaluate soil behavioral type (SBT) that uses normalized piezocone readings (Robertson 1990
1991 Lunne et al 1997)
A-10
(119902119905 minus 120590119907119900) (a) normalized tip resistance 119876 = A11 prime 120590119907119900
100∙119891119904 (b) normalized sleeve friction 119865() = A12
(119902119905 minus 120590119907119900)
(1199062 minus 1199060) (c) normalized porewater pressure 119861119902 = A13
(119902119905 minus 120590119907119900)
While the Q-F-Bq classification is a three-dimensional plot it is often presented as a set of two-dimensional
plots specifically Q vs F and Q vs Bq as shown in Figure A6 Data are grouped according to layers and
can be superimposed to identify their association with the nine soil behavioral types All indirect CPT soil
classification approaches should be cross-checked and verified for a particular geologic setting and local
geotechnical conditions before routine use in practice
Figure A6 Nine-zone soil behavioral type charts (a) normalized cone resistance vs normalized
sleeve friction (b) normalized cone resistance vs normalized porewater pressure parameters
(adapted after Robertson 1991 Lunne et al 1997)
The development of a CPT material index (Ic) has been found advantageous in the initial screening of soil
types and helps to organize the sounding into their respective SBT zones of similar soil response In this
case the CPT index is found from (Robertson 2009)
A-11
A2 22 )log221()log473( rtnc FQI
where Qtn = modfied stress-normalized net cone resistance given by
(119902119905 minus 120590119907119900)120590119886119905119898 119876 = A31 prime (120590119907119900120590119886119905119898)119899
where σatm = reference stress equal to atmospheric pressure (1 atm asymp 1 bar = 100 kPa asymp 1 tsf asymp 147 psi)
The exponent n is soil-type dependent n = 1 (clays) n asymp 075 (silts) and n asymp 05 (sands) If units of bars are
used then Qtn (units of bars) is simply
(119902119905 minus 120590119907119900) 119876 = 119899 A32
prime ) (120590119907119900
The operational value of exponent n is found by iteration Initially an exponent n = 1 is used to calculate
the starting value of Ic (ie Qtn = Q) and then the exponent is upgraded to
prime 120590119907119900 119899 = 0381 ∙ 119868119888 + 005 ( ) minus 015 le 10 A4 120590119886119905119898
Then the index Ic is recalculated Iteration converges quickly and generally only 3 cycles are needed to
secure the operational Ic at each depth The soil zones and associated Ic values are detailed in Figure A7
The sensitive soils of zone 1 can be screened using the following expression
A-12
Zone 1 119876119905119899 lt 12119890119909119901( minus14 ∙ 119865119903)
A5
If any soil layers are found within zone 1 (sensitive soils) then caution should be exercised as these
structured clays are prone to instability collapses and difficulties in construction and performance
Another indicator of zone 1 sensitive soils is when the normalized porewater parameter Bq gt 08 When
these criteria are evident the geoengineer should consult with senior geotechnical personnel or the chief
engineer for guidance
Figure A7 Delineation of soil behavioral type zones using CPT material index Ic
A-13
Stiff overconsolidated clayey sands and sandy clays soils of zone 8 (15 lt Fr lt 45) and zone 9 (Fr gt
45) can be identified from the following criterion
Zones 8 and 9 0020)1(00030)1(0050
12
rr
tnFF
Q A6
The remaining soil types are identified by the CPT material index Zone 2 (organic clayey soils Ic ge 360)
Zone 3 (clays to silty clays 295 le Ic lt 360) Zone 4 (silt mixtures 260 le Ic lt 295) Zone 5 (sand mixtures
205 le Ic lt 260) Zone 6 (clean sands 131 le Ic lt 205) and Zone 7 (gravelly to dense sands Ic le 131)
The red dashed line at Ic = 260 represents an approximate boundary separating drained (Ic lt 260) from
undrained behavior (Ic gt 260)
Once the specific zone has been assigned to a layer a visual representation can be made to show
either the zone number or colorization so that the predominant layers and soil types can be
realized
A123 Probabilistic Soil Types from CPT
The inference of soil type has been addressed using probabilistic methods as discussed by Abu-
Farshakh et al (2008) and Tumay et al (2013) Here the CPT readings can be post-processed to
provide the percentage of sand silt and clay components with depth The output is consistent
and compatible with the current MnDOT soil classification chart (Figure A8) which is similar in
content to that used by the USDA
A-14
Figure A8 Chart for soil classification system by MnDOT
A13 Identifying Soft Organic Soils
The identification of soil type using cone penetration tests (CPT) is usually done on the basis of empirical
soil behavioral type (SBT) charts that use the measured readings (qt fs andor u2) While there at least
25 different sets of charts available (ie Hegazy 1998 Eslami et al 2017 Valsson 2016) one of the most
widely used and popular is the 9-zone SBT system that uses three normalized cone readings Qtn Fr and
Bq (Robertson amp Cabal 2007 Lunne et al 1997) The system is denoted as SBTn and plotted on charts in
terms of Q vs F and Q vs Bq as shown in Figure A9
A-15
1
10
100
1000
01 1 10
Norm
aliz
ed
Con
e T
ip R
esis
tan
ce
Q
Normalized Friction Fr = 100 fst(qt - svo) ()
9-Zone Soil Behavioral Type Chart
Gravelly Sands(zone 7)
Sands
(zone 6)
Sandy Mixtures(zone 5)
Silt Mix(zone 4)
Clays
(zone 3)
Organic Soils(zone 2)
Sensitive Clays
and Silts(zone 1)
Very stiff OC clayto silt (zone 9)
Very stiffOC sandto clayey
sand(zone 8)
1
10
100
1000
-06 -04 -02 00 02 04 06 08 10 12 14
No
rmal
ized
Co
ne
Tip
Res
ista
nce
Q
Normalized Porewater Bq = Du2(qt-svo)
Clays(zone 3)
Sensitive(zone 1)Organic
(zone 2)
Gravelly Sands(zone 7)
Sands(zone 6)
Figure A9 CPT charts for SBTn system (a) Q versus F (b) Q versus Bq
Of particular concern in geotechnical site characterization is the proper identification of soft
organic soils such as organic clays and silts (OL and OH) and peats (Pt) These soils often cause
difficulties and problems during construction and performance of civil engineering works because
of bio-degradation long-term creep excessive settlements andor other issues
The SBTn system has a category labelled Zone 2 recognizes organic soils However several recent
studies have noted that data from CPTu soundings do not always fall within the bounds
delineated using Zone 2 even though laboratory tests and field inspections clearly note the
presence of organic soils Lab tests to classify organic soils includes loss on ignition (LOI gt 5)
and two pairs of Atterberg limits testings per ASTM standards as well as the visual-manual
identification of strong organic odor and smell and dark color notably black and dark gray
Results by Mlynarek et al (2014) show that organic soils in Poland are not adequately identified by the
Q-F chart as seen in Figure A10 Moreover studies by Zawrzykraj et al (2017) found that neither the Q-
A-16
F and Q-Bq plots worked well to recognize organic soils and peats in Poland Nejaim et al (2016) found
that Brazilian soft organic clays were misclassified as Zone 3 (clays) rather than Zone 2 (organic clays)
Similarly a recent study on CPTu data from organic clays located at 24 sites in the USA Sweden Mexico
Brazil and Australia have been compiled (Agaiby 2018) These data are shown to generally avoid falling
with the Zone 2 boundaries in either Q-F or Q-Bq charts thus are not recognized during these soundings
as indicated by Figure A11 The exception in this case is the Mexico City clay which is correctly identified
however the other 23 sites are generally not properly recognized and categorized Additional findings by
Missiaen et al (2015) found that the CPTu results in Belgian soils also did not identify organic clays and
peats in the non-normalized version of these charts that uses Qt versus friction ratio (Rf = 100middotfsqt) as
detailed by Robertson amp Cabal (2007)
Figure A10 Soil behavioral chart with superimposed CPTu data from Polish organic soils
(from Mlynarek et al 2014)
A-17
Figure A11 Paired SBTn charts with CPTu data from 24 organic clay sites indicating non-compliance
with Zone 2 (organic soils)
(compiled by Agaiby 2018)
A14 Simplified Yield Stress Evaluation in Clays by CPTu
From the development of a hybrid formulation of spherical cavity expansion (SCE) theory with critical-
state soil mechanics (CSSM) the effective preconsolidation stress or yield stress (σp) of clays can be
expressed in terms of three piezocone parameters (a) net cone tip resistance qnet = qt - σvo (b)
measured excess porewater pressure Δu2 = u2 - u0 and (c) effective cone resistance qE = qt - u2 Details
on the SCE-CSSM solutions are given elsewhere (Mayne 1991 Mayne 2017 Agaiby amp Mayne 2018)
A-18
For normal and well-behaved clays which include inorganic fine-grained soils such as clays and silts
of low sensitivity a set of simple relationships can be derived By adopting characteristic default values
for effective friction angle ϕ = 30deg and rigidity index (IR = 100) linear expressions for the effective
preconsolidation stress are obtained
prime 120590119901 = A7 033 ∙ 119902119899119890119905
prime 120590119901 = 053 ∙ ∆1199062 A8
prime 120590119901 = 060 ∙ 119902119864 A9
A141 Application to soft Chicago clay
An example of a common geomaterial in this category includes the infamous soft Chicago clays
that were deposited in a glacio-lacustrine environment (Cho amp Finno 2010) The soft clay has a
mean water content of 20 liquid limit of 38 and plasticity index PI= 12 Results of CPTu tests
conducted at the national geotechnical experimentation site (NGES) at Northwestern University
are shown in Figure A12 (Ouyang amp Mayne 2017) As seen in the profile a thin soft clay resides
between depths of 9 to 105 m with a thicker soft clay layer in the range of 12 to 18 m
A-19
0
2
4
6
8
10
12
14
16
18
20
00 02 04 06 08 10 12D
ep
th (
m)
CPTu qt and u2 (kPa)
sand (SP)
Blodgett
Park Ridge
Deerfieldsoft clay
ClayLayerof Study
Northwestern UnivNational Test Site
Figure A12 Results of CPTu sounding at the NGES at Northwestern University Illinois
Post-processing of the CPTu data using Equations A7 A8 and A9 show good agreement in evaluating the
profile of effective yield stress in this clay layer compared with results from one-dimensional
consolidation tests made on undisturbed samples taken at the site
A-20
10
11
12
13
14
15
16
17
18
19
20
0 100 200 300 400 500 600 700 800 900 1000
De
pth
(m
)Effective Yield Stress sp (kPa)
Consols
033 qnet
053 Du2
060 qE
svo
Figure A13 Evaluation of effective yield stresses at NWU using consolidation tests and CPTu
A142 Application to soft Bay Mud California
Another example of a well-behaved fine-grained soil is the San Francisco Bay Mud which is a soft clay
deposit in northern California Results of a representative CPTu are shown in Figure A14 and indicate the
soil profile at a site in eastern San Francisco (Hunt et al 2002) For this site index values include 70 lt
wn lt 90 60 lt LL lt 80 and 35 lt PI lt 45 Post-processing the CPTu data using Equations A7 A8 and
A9 show good agreement with each other as well as with the results of consolidation tests as seen in
A-21
Figure A15 The consolidation tests were performed using a CRS device as reported by Pestana et al
(2002)
San Francisco Bay Mud (Pestana et al 2002)
0
5
10
15
20
25
30
35
40
0 1000 2000 3000
De
pth
(m
ete
rs)
Cone Resistance qt (kPa)
0
5
10
15
20
25
30
35
40
0 10 20 30 40 50 60
Sleeve Friction fs (kPa)
0
5
10
15
20
25
30
35
40
0 500 1000 1500
Porewater u2 (kPa)
Miscellaneous Fill
Young Bay Mud
Young Bay Mud
Clayey Sand
Miscellaneous Fill
Young Bay Mud
Young Bay Mud
Figure A14 Representative CPTu in San Francisco Bay Mud (data from Hunt et al 2002)
A-22
San Francisco Bay Mud (Pestana et al 2002)
0
5
10
15
20
25
30
0 100 200 300 400 500 600
De
pth
(m
ete
rs)
Preconsolidation Stress sp (kPa)
svo
033 qnet
053 Du2
060 qE
CRS
svo
Figure A15 Evaluation of effective yield stresses in Bay Mud from consolidation tests and CPTu
A15 CPTu Screening of Soft Organic Clays
For soft organic clays the Equations A7 A8 and A8 will not apply thus can serve as a means to
screen such geomaterials from the soil profile Two examples of CPTu soundings in soft organic
clays are presented to illustrate the approach Additional case records and documentation of
CPTu data in organic clays are given in Agaiby (2018)
A151 Soft organic alluvial soils in Washington DC
Results of in-situ tests including CPTu soundings in soft organic clayey silts along the Potomac River at the
Anacostia Naval Air Station are presented by Mayne (1987) Here the soft soils are categorized as organic
clayey silts (OH) per the Unified Soils Classification Systems (USCS) with mean values (and plus and minus
A-23
one standard deviations) of wn = 68 plusmn 16 LL = 83 plusmn 25 and PI = 37 plusmn 17 A representative
piezocone sounding is depicted in Figure A16 For the post-processing of CPTu data the use of the Q-F
and Q-Bq graphs do not find any points within the Zone 2 category for organic soils When the data are
evaluated to determine the profile of effective yield stress the Equations A7 A8 and A9 do not agree as
shown by Figure A17
0
5
10
15
20
25
0 2000 4000 6000
Dep
th (
met
ers)
Cone Resistance qt (kPa)
0
5
10
15
20
25
0 20 40 60 80 100
Sleeve Friction fs (kPa)
0
5
10
15
20
25
0 200 400 600
Pore Pressure u2 (kPa)
Figure A16 Representative CPTu in soft organic clay along the Potomac River DC
A-24
0
5
10
15
20
25
0 3 6 9D
epth
(m
eter
s)Soil Behaviour Type Zone
Q and Fchart
0
5
10
15
20
25
0 100 200 300 400 500 600
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
033qnet
053 Du2
06 qE
Oed
Figure A17 Post-processing of CPTu data from Anacostia-Bolling (a) SBTn zones using CPT material
index Ic (b) CPTu screening equations for yield stress
Note that the hierarchy of the results from Equations A7 A8 and A9 show that the CPTu
evaluations for preconsolidation stress in soft organic clays indicates
[A10] σp = 053 Δu2 lt 033 qnet lt 060 qE
A152 Soft organic peaty clay in Saint Paul MN
Results of CPTu tests were collected during a training exercise underneath I-35E just northeast
of Saint Paul MN in 2007 All three MnDOT CPT rigs available at the time were used to obtain
in-situ test data at the site The site is well known as having soft organic soils and samples were
obtained from three borings made at the site (Boring ID T523 T542 and T518) Boring logs
A-25
indicate the presence of highly organic silt loams with marl and spongy organic clayey silts
From 12 samples the natural water contents ranged from 30 to 191 with a mean of 112 plusmn
58
A representative piezocone sounding from the site (MnDOT CPTu ID F22Y0703C) is used for
this illustration and is presented as Figure A18 Post-processing these data using the SBTn
algorithms and CPT material index show that the soils classify primarily as Zone 3 (clays) and
Zone 4 (silty mix) thus do not identify the geomaterials properly as Zone 2 (organic soils)
0
5
10
15
0 500 1000 1500 2000
Dep
th (
met
ers)
Cone Tip Resistance
qt (kPa)
0
5
10
15
0 10 20 30 40 50 60
Sleeve Friction
fs (kPa)
0
5
10
15
-100 0 100 200 300 400
Porewater Pressure
u2 (kPa)
Figure A18 MnDOT CPTu sounding at the I-35E test site near St Paul Minnesota
A-26
1
10
100
1000
01 1 10
Norm
aliz
ed
Con
e T
ip R
esis
tan
ce
Q
Normalized Friction Fr = 100 fst(qt - svo) ()
9-Zone Soil Behavioral Type Chart
St PaulGravelly Sands(zone 7)
Sands
(zone 6)
Sandy Mixtures(zone 5)
Silt Mix(zone 4)
Clays
(zone 3)
Organic Soils(zone 2)
Sensitive Clays
and Silts(zone 1)
Very stiff OC clayto silt (zone 9)
Very stiffOC sandto clayey
sand(zone 8)
1
10
100
1000
-06 -04 -02 00 02 04 06 08 10 12 14
No
rmal
ized
Co
ne
Tip
Res
ista
nce
Q
Normalized Porewater Bq = Du2(qt-svo)
Clays(zone 3)
Sensitive(zone 1)Organic
(zone 2)
Gravelly Sands(zone 7)
Sands(zone 6)
Figure A19 Post-processing St Paul sounding using the 9-zone SBTn classification system
Using the recommended approach Figure A20 shows the CPT-evaluated yield stress profiles from
equations A7 A8 and A9 Here again the three estimates of σp do not agree but show the same
hierarchy as detailed in Equation A10
A-27
0
5
10
15
20
0 50 100 150 200 250 300 350 400
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
06 qeff svo
033qnet Consols
054 Du2
Figure A20 CPTu screening of soft organic soils at St Paul test site
A16 CPTu Evaluations of Yield Stress in Soft Organic Soils
Once the presence of soft organic clays and silts is identified using the aforementioned CPT
screening algorithms the evaluation of effective yield stress is conducted using the procedure
outlined in Chapter 2 of this MnDOT manual For soft organic soils an exponent of m = 090 is
recommended (Mayne et al 2009) Therefore the preconsolidation stress is obtained from
(units of kPa)
prime = A11 120590119901 033 ∙ (119902119899119890119905)090
The algorithm can be expressed in dimensionless form by
prime = A12 120590119901 033 ∙ (119902119905 minus 120590119907119900)119898prime ∙ (120590119886119905119898100)1minus119898prime
A-28
where σatm = reference stress (= 100 kPa = 147 psi) and m = 090 for soft organic silts and clays
The approach is applied to the two former example case studies in Figure A21 (Washington DC)
and Figure A22 (St Paul MN) respectively showing good agreement with benchmark results
taken from laboratory consolidation tests on undisturbed samples of these sites in both cases
0
5
10
15
20
25
0 100 200 300 400 500 600
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
033qnet^090
Oed
Figure A21 Profiles of yield stress from consolidation tests and CPTu method for organic clayey silts at
Anacostia-Bolling site in Washington DC
A-29
0
5
10
15
20
0 20 40 60 80 100 120 140 160
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
CPT
Consols
120590119901prime = 033 119902119899119890119905
090 in kPa
Figure A22 Profiles of yield stress from consolidation tests and CPTu method for organic clayey silts at
MnDOT test site near Saint Paul Minnesota
A17 Soil Unit Weight
As shown by Figure A23 total soil unit weight (t) can be estimated from the CPT sleeve friction resistance
(Mayne 2014) The equation can be expressed by
120574119905 = 120574119908 ∙ [122 + 015 ∙ 119897119899 (100 lowast 119891119904120590119886119905119898 + 001)] A13
where w = unit weight of water and satm = atmospheric pressure
A-30
Figure A23 Evaluation of soil unit weight from CPT sleeve friction
A18 Effective Stress Friction Angle
Sands
The effective stress friction angle () is one of the most important soil properties as it governs the strength
of geomaterials as well as affects soil-pile interface and pile side friction While an effective cohesion
intercept (c) can also be considered this is usually reserved for cemented or bonded geomaterials or
unsaturated soils and may lose its magnitude with time ageing or with prolonged environmental
changes
For clean quartz to silica sands where porewater pressures are essentially hydrostatic (Bq = 0) the
following expression has been calibrated with triaxial compression test results from undisturbed sand
samples and normalized cone resistances as presented in Figure A24 (Mayne 2007 2014)
A-31
120601prime(119889119890119892) = 176deg + 110deg 119897119900119892 (1199021199051) A14
Figure A24 Evaluation of effective stress friction angle in quartz-silica sands from CPT
Note Relationship applies to drained soil behavior when Ic lt 26 andor Bq lt 01
where qt1 is an earlier form of stress normalized cone tip resistance for sands given by
(119902119905 120590119886119905119898) 1199021199051 = A151 05 prime (120590119907119900120590119886119905119898)
A-32
In terms of units of bars this is expressed simply as (in units of bars)
119902119905 1199021199051 = 05 A152 prime ) (120590119907119900
Recently Robertson amp Cabal (2015) recommended the use of the modified form of normalized cone
resistance (Qtn) in Equation A10
120601prime (119889119890119892) = 176deg + 110deg 119897119900119892 (119876119905119899) A16
In sands qnet asymp qt since the overburden term is small relative to the cone tip resistance Figure A25 shows
that the two normalizations give very comparable results
Figure A25 Comparable values from qt1 and Qtn normalization schemes for CPTs in sands
Clays and Silts
A-33
In the case of soft to firm intact clays and silty clays the effective friction angles are determined from the
normalized cone resistance and porewater pressure parameters (Senneset et al 1989 Mayne 2016) as
shown in Figure A26 The exact solution when the angle of plastification = 0 is given as
qBQ
)tan1(tan61
1)tanexp()245(tan 2
A17
Approximate NTH Solution for from CPTu
qBQ
)tan1(tan61
1)tanexp()245(tan2
Approx asymp 295deg∙Bq0121∙[0256 + 0336∙Bq + log Q ]
1
10
100
20 25 30 35 40 45
Co
ne
Res
ista
nce
Nu
mb
er Q
(degrees)
Bq
010204060810
c = 0 = 0deg
Theory (dots)
Approximation (lines)
Theory
Figure A26 Evaluation of effective stress friction angle in clays and silts from CPTu results
Note Relationship applies to undrained soil behavior when Ic ge 26 andor Bq ge 01
which can be approximately inverted into the form (Mayne 2007)
A-34
0121 120601prime = 295deg ∙ 119861119902 ∙ [ 0256 + 0336 ∙ 119861119902 + log(119876)] A18
This algorithm is specifically applicable for the following ranges of porewater pressure parameter (01 le
Bq lt 1) and effective stress friction angles (20deg le lt 45deg)
A19 Stress History
The stress history can be characterized by an apparent yield stress or preconsolidation stress (sp) as well
as by its normalized and dimensionless form YSR = spsvo = yield stress ratio The YSR is in effect the
same as overconsolidation ratio (OCR) however now generalized to accommodate mechanisms of
preconsolidation that occur beyond just erosion glaciation and removal of overburden stresses but also
due to ageing desiccation repeated cycles of wetting-drying bonding repeated freeze-thaw cycles
groundwater changes and other factors
A-35
A generalized approach for evaluating the yield stress or preconsolidation in soils using net cone
resistance has been formulated as presented in Figure A27 (Mayne 2015) Yield stress in soils from CPT
10
100
1000
10000
10 100 1000 10000 100000
Yie
ld S
tre
ss
s
y
(kP
a)
Net Cone Resistance qt - svo (kPa)
Blessington
General Trend
sy = 033(qt-svo)m
Fissured Clays m = 11
Intact clays m = 10 Sensitive clays m = 09
Silt Mixtures m = 085Silty Sands m = 080
Clean Sands m = 072
m = 11 10 09
085
080
072
Units kPa
Fissured Clays
Figure A27 Evaluation of yield stress or preconsolidation stress in soils from CPT
The algorithm can be expressed in dimensionless form by
1minus119898prime prime 120590119886119905119898 120590119901 = 033 ∙ (119902119905 minus 120590119907119900)119898prime
( ) A19 100
where m = exponent depends on soil type m = 1 (intact clays) 085 (silts) 080 (sandy silts to silty sands)
and 072 (sands) For fissured clays the exponent m may be 11 or higher depending upon the age of the
A-36
formation and degree of jointing and discontinuities Note that fissured clays can be identified when Ic lt
26 and Bq lt 01
The exponent m has also been calibrated with CPT material index as presented in Figure A28
An algorithm to express this relationship for non-fissured soils is given by
25)652(1
2801
cIm
A20
This allows the post-processing of CPT to automatically choose the appropriate exponent m for
a layer by layer analysis
Figure A28 Yield stress exponent m in terms of CPT material index Ic
A-37
A110 Lateral Stress Coefficient
The horizontal geostatic state of stress is represented by the lateral stress coefficient K0 = shosvo
commonly referred to as the at-rest condition The magnitude of K0 for soils that have been loaded and
unloaded can be approximately estimated from
119870119900 = (1 minus 119904119894119899120601prime) ∙ 119874119862119877119904119894119899120601prime A21
Data compiled from in-situ K0 measurements using self-boring pressuremeter tests (SBPMT) total stress
cells (TSC) or push-in spade cells andor laboratory methods (instrumented consolidometers triaxials
andor suction measurements) on a variety of clays silts and sands have been compiled and reported by
Ku amp Mayne (2015) as shown in Figure A29 verifying Equation A15 as a means for evaluating K0 in soils
Figure A29 Relationship between lateral stress coefficient K0 and YSR or OCR in soils (Ku amp Mayne
2015)
A-38
A111 Undrained Shear Strength
The loading of soils can occur under conditions of being fully drained (Du = 0) partially drained or full
undrained (DVV0) where Du = excess porewater pressures (above hydrostatic) and DVV0 = volumetric
strain Further details on specific stress paths are best explained in terms of critical state soil mechanics
or CSSM (Mayne et al 2009 Holtz Kovacs amp Sheahan 2011) The prevailing drainage conditions depend
upon the rate of loading and permeability characteristics of the soil Normally in sands that are pervious
and exhibit high permeability a drained response occurs Exception to this may occur in loose sands
during fast earthquake loading resulting in soil liquefaction In clays that exhibit low permeability a fast
rate of loading will result in undrained loading at constant volume This in fact is a temporary and
transient condition often termed short term loading In the long term eventually porewater pressures
will dissipate (albeit slowly) and a drained response will prevail thus termed long term loading
The overall soil strength is controlled by the effective stress strength envelope Most commonly this is
represented by a simple linear relationship termed the Mohr-Coulomb criterion where the maximum
shear stress (max called the shear strength) is given as
= 119888prime + 120590prime ∙ 119905119886119899 120601prime A22 120591119898119886119909
where c = effective cohesion intercept s = effective normal stress and = effective stress friction angle
As a starting point values of c = 0 and = 30deg can be adopted for all soil types at least until the specific
soil type geologic formation and results from high-quality laboratory or field test data are available
Peak Undrained Shear Strength from Stress History
Using simplified critical state soil mechanics the peak undrained shear strength (max = su) can be
evaluated from the effective stress strength envelope (c = 0 effective friction angle ) and stress history
(ie OCR) in the form
prime DSS 119904119906 = (119904119894119899120601prime2) ∙ (119874119862119877)Λ ∙ 120590119907119900 A23
A-39
which corresponds to a simple shear mode Figure A30 presents the aforementioned relationship
together with data from 17 different clays tested in simple shear
Figure A30 Normalized undrained shear strength from simple shear tests on various clays
versus YSR or OCR
For the triaxial compression (TC) mode the equation would give slightly higher strengths that are
calculated from
prime TC 119904119906 = (1198721198882) ∙ (1198741198621198772)Λ ∙ 120590119907119900 A24
where Mc = (6 middot sin)(3-sin)
Peak Undrained Shear Strength from CPTu
A more direct approach to assessing undrained shear strength is via bearing capacity theory
whereby
A-40
A25 kt
votu
N
qs
s
where Nkt = bearing factor that depends on mode of shearing sensitivity of the clay degree of
overconsolidation and other factors For soft offshore clays the back figured Nkt from data collected at
14 well-documented sites (Low et al 2010) determined mean values based on mode of shearing Nkt =
119 (triaxial compression) Nkt = 136 (simple shear) and Nkt = 133 (vane)
A recent study of 51 clays that were tested by both field piezocone and laboratory CAUC triaxial tests
showed that essentially Nkt = 12 for intact clays and clayey silts of low to medium sensitivity (Mayne
Peuchen amp Baltoukas 2015) Figure A31 shows the slightly different trends for offshore versus onshore
clays For sensitive clays a lower value of Nkt = 10 would be appropriate and for fissured clays a higher
value of Nkt would be in the range of 20 to 30
A-41
A-42
Figure A31 Database from 51 clays with CPT qnet versus lab measured triaxial shear strength
(Mayne Peuchen amp Baltoukas 2015)
For soft to firm clays an alternate means to evaluate undrained shear strength is via the excess porewater
pressures ( u = u2 - u0) from the following
u
uN
us
D
D 2
A26
where N u = porewater bearing factor also dependent on the aforementioned factors The study by
Low et al (2010) found N u = 59 (triaxial compression) N u = 69 (simple shear) and N u = 71 (vane
shear)
A third alternate is found from the effective cone resistance (qE = qt - u2) whereby
kE
tu
N
uqs 2
A27
where NkE is a bearing term for this approach Mayne et al (2015) indicated a mean value of NkE = 8 for
soft-firm intact clays
Remolded Undrained Shear Strength from CPT
The remolded undrained shear strength (sur) is obtained from either field vane tests lab mini-vane shear
tests or lab fall cone devices This affords the evaluation of the clay sensitivity (St) which is defined as the
ratio of peak to remolded strengths at a given water content
119904119906(119901119890119886119896) 119878119905 = frasl A28 119904119906119903
For the CPT the sleeve friction has been noted to give values that are comparable to the
remolded undrained shear strength (Powell amp Lunne 2015) Therefore
asymp 119891119904 A29 119904119906119903
A112 Ground Stiffness and Soil Moduli
The stiffness of the ground can be represented by several geoparameters including the compressibility
indices (Cr Cc Cs) spring constants (kv) and moduli Regarding the latter there are several moduli that
are used in geotechnical engineering including shear modulus (G and Gu) Youngs modulus (E and Eu)
constrained modulus (D) bulk modulus (K) resilient modulus (MR) and subgrade reaction modulus (ks)
A-43
Soil Modulus
The definition of modulus is taken as E = DsD with Figure A32 showing a typical deviator stress (q = s1
- s3) versus axial strain (a) curve Theoretical interrelationships between the elastic moduli G E D and
K have a dependence on the Poissons ratio v The value of Poissons ratio can be taken as vu = 05 for
undrained loading (ie constant volume) while for drained loading which is accompanied by volumetric
strains a value of v = 02 may be used If we adopt the reference modulus as E then the
interrelationships with the other elastic moduli are given by
a Reference Stiffness 119864prime = drained Youngs modulus A30
119864prime
b Shear Modulus 119866prime = A31 [2(1+120584prime)]
119864prime (1minus120584prime) c Constrained Modulus 119863prime = A32
[(1+120584prime)(1minus2120584prime)]
119864prime
d Bulk Modulus 119870prime = A33 [3middot(1minus2120584prime)]
A-44
Figure A32 Representative stress-strain-strength curve for soils in triaxial compression
For the extreme case when = 0 in fact D = E When = 02 the two moduli are only 10
different D = 11middotE thus the constrained modulus and drained Youngs modulus are often
considered somewhat interchangeable
The resilient modulus (MR) applies to pavement analysis and design most commonly measured by cyclic
triaxial testing under repeated load applications In fact MR is a special case of Youngs modulus that
relates to the small strain stiffness measured in the nondestructive range but has developed permanent
plastic strains after many cycles of loading (Brown 1996 Dehler amp Labuz 2007)
The subgrade reaction modulus is actually a combined soil-structural parameter as its value
depends on the ground stiffness and the size of the loaded element The subgrade modulus is
defined as ks = q where q = applied stress and = measured deflection In terms of elasticity
solutions the deflection of a flexible circle of diameter d is given by = qmiddotdmiddot(1-2)E Therefore
119864prime
119896119904 = A34 [119889middot(1minus1205842)]
which has units of kNm3 or pcf
Table A2 summarizes several selected studies towards the approximate evaluation of these moduli
obtained directly from measured CPT data Note however that the results from in-situ penetrometer
tests generally represent a peak strength as indicated in Figure A32 Thus the measured cone tip
resistance (qt) reflects the top of the stress-strain curve either the undrained strength in clays (su) or the
effective friction angle () of sands or even a strength intermediate between these two values Thus
A-45
Source Soils Studied Reference Modulus
Findings
Kulhawy amp Mayne 1990
8 stiff OC clays Lab constrained modulus
D asymp 825 middot (qt - svo)
Mohammed et al (2000)
Resilient modulus
Abu-Farsakh 2004
7 Louisiana sites
Lab constrained modulus
D asymp 358 middot (qt - svo)
Mayne (2007b)
All soil types Constrained moduli from lab consolidation
First-order estimate D asymp middot(qt - svo)
Robertson (2009)
All soil types Constrained modulus from consolidation
D = m middot (qt - svo)
when Ic gt 22
1 m = Qtn when Qtn le 14
2 m = 14 when Qtn gt 14
when Ic lt 22 then
= 003 middot 10055 Ic + 168 m
Liu et al (2016)
16 clay sites in China
Resilient modulus MR = (146qt 053 + 1355 fs
14 + 236)244
(all in MPa)
Casey et al (2016)
8 clays Triaxial tests for undrained Youngs modulus
Eu(svc)07 = 316 - 23 middot LL
where Eu and svc (MPa)
and LL = liquid limit ()
qt is probably not the best means to obtain the slope of the stress-strain curve Instead the SCPTu that
provides the initial tangent shear modulus (Gmax) from the shear wave velocity (Vs) is a better choice
Table A2 Selected Modulus Relationships with Cone Penetration Test Measurements
A-46
Nonlinear Modulus
The small-strain shear modulus (Gmax or G0) represents the initial stiffness of all soils and rocks It is the
beginning of all stress-strain-strength curves for geomaterials and obtained from elasticity theory
119866119898119886119909 = 1198660 = 120588119905 middot 1198811199042 A35
where Vs = shear wave velocity t = tga = total soil mass density t = total soil unit weight and ga =
gravitational acceleration constant (98 ms2)
From a general viewpoint on stiffness the shear modulus G of soil is defined as the slope of shear stress
versus shear strain G = DDs for a tangent definition and by G = s as a secant definition The shear
modulus is related to its associated Youngs modulus E = 2G(1+v) Both moduli are in fact highly
nonlinear ranging from a maximum value at the small-strain stiffness (Gmax = G0 = t middotVs2 where t =
total soil mass density and Vs = shear wave velocity) to intermediate G values at medium strains (asymp 1)
to low values at peak strength As such a variety of algorithms and formulae have been developed to
represent either a partial range or the full stress-strain-strength behavior of soils over a range of
interests (eg Mayne 2005) In these formulations a variable number of input parameters may be
required in order to produce a stress-strain-strength curve
A modified hyperbolic form suggested by Fahey amp Carter (1993) has favorable attributes in that the
modulus reduction factor can be established with only a single variable This allows the initial stiffness
to the be small-strain stiffness (G0) that is reduced in terms of level of mobilized strength eg max =
qqmax = 1FS which is simply the reciprocal of the calculated factor of safety (FS) The magnitude of
secant shear modulus (G) corresponding to the particular level of loading is given by
119866 = 119872119877119865 middot 1198660 = (1198661198660) middot 1198660 A36
A-47
where the MRF = modulus reduction factor determined as
1 minus (120591 )119892 119872119877119865 = (1198661198660) = frasl A37 120591119898119886119909
and g = empirical exponent term Figure A33 shows the relationships for normalized shear stress ()
versus shear strain (s) and corresponding normalized secant shear modulus (G = s) with shear strain
over a range of exponent values 01 le g le 10 For a discussion of tangent G please see Fahey amp Carter
(1993)
Using laboratory data from resonant column-torsional shear tests and triaxial specimens with local
strain measurements for Gmax reference values modulus reduction curves for a selection of sands and
clays are presented in Figure A34 The data indicate that the exponent g falls within the range 02 lt g lt
05 for many soils tested drained and undrained A mean value of g = 03 is recommended for
preliminary site investigations and designs until additional information can be obtained
The value of max is the shear strength of the soil generally taken as either (1) drained (c = 0) max =
svo middot tan or (2) undrained where max = su = undrained shear strength as discussed earlier Thus each
shear stress () can be associated with its shear modulus (G) and the relevant shear strain is found from
s = G This allows for generation of nonlinear stress-strain-strength curves at all depths from SCPTu
data in clays sands and mixed soil types
A-48
Modified Hyperbola (Fahey amp Carter 1993 Canadian Geotech J) Coefficient Term f = 1
``
0
01
02
03
04
05
06
07
08
09
1
000001 00001 0001 001 01 1
No
rmal
ized
GG
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
0
01
02
03
04
05
06
07
08
09
1
0 01 02 03 04 05 06 07 08 09 1
No
rmali
zed
GG
ma
x
Mobilized Shear Stress max
g =1
07
05
03
02
01
0
01
02
03
04
05
06
07
08
09
1
00001 0001 001 01 1
No
rmali
zed
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
0
01
02
03
04
05
06
07
08
09
1
0 1 2 3 4 5 6 7 8 9 10
No
rmali
zed
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
Reduction Factor
GGmax = 1- (max)g
Figure A33 Normalized modulus (GGmax) and normalized shear stress (max) versus shear strain (s)
for modified hyperbolic relationship of Fahey amp Carter (1993)
A-49
0
01
02
03
04
05
06
07
08
09
1
0 01 02 03 04 05 06 07 08 09 1
Mo
du
lus
Re
du
cti
on
G
Gm
ax
or
EE
ma
x
Mobilized Strength max or qqmax
NC SLB Sand
OC SLB Sand
Hamaoka Sand (e=0625)
Hamaoka Sand
Toyoura Sand e = 067
Toyoura Sand e = 083
Ham River Sand
Ticino Sand
Kentucky Clayey Sand
Kaolin
Fujinomori Clay
Pietrafitta Clay
London Clay
Vallericca Clay
Pisa Clay
Pisa Clay
Pisa Clay
Pisa Clay
Kiyohoro Silty Clay
Thanet Clay
Exp g = 02
Exp g = 03
Exp g = 04
Exp g = 05
MRF = 1 - (qqmax)g
Figure A34 Modulus Reduction Factors (MRF) with mobilized strength of clays and sands
A113 Coefficient of Consolidation
The rate at which foundation and embankment settlements occur as well as the dissipation of excess
porewater pressures is controlled by the coefficient of consolidation (cv) The magnitude of cv is also
required in the design of vertical wick drains that can be installed in soft ground to expedite the time for
consolidation Using the results of CPTu dissipation tests which measure the rate at which the u2 readings
vary with time the in-situ profile of cv can be evaluated Often the piezocone-interpreted values of cv
are validated by comparison with results from laboratory one-dimensional consolidation tests (eg Abu-
Farsakh MY 2004) Alternatively a better method is to cross-check the values with the measured full-
scale performance of instrumented embankments that are constructed over the ground and the recorded
time-rate of consolidation and settlements can provide the best cv for that site (Abu-Farsakh MY et al
2011)
A-50
Monotonic Dissipation Tests
An illustrative dissipation curve from piezocone testing at IDT State Highway 95 at an embankment and
bridge crossing is presented in Figure A35 After the CPTu sounding was halted at a depth of 512 feet
the recorded decay of porewater pressures was observed to be monotonic with time until the dissipations
were ended at 1000 seconds During that time the u2 readings decreased from 115 psi to 47 psi At this
location the depth to the groundwater table is about 16 feet thus the calculated equilibrium water
pressure is 15 psi Commonly a characteristic time for piezo-dissipations is taken at 50 degree of
consolidation although other degrees may be adopted By evaluating the value of u2 at 50 as shown in
Figure A35 the characteristic t50 = 404 s can be obtained
0
20
40
60
80
100
120
1 10 100 1000 10000
Pore
wat
er P
ress
ure
u2
(psi
)
Time (s)
CPTU-02E-1 z = 1560 m
uo (hydrostatic)
u2 (initial) = 115 psi
u2 (final projected) = u0 = 15 psi
u2 at 50 = frac12(115+15) = 65 psi
Idaho DOT State Route 95Dissipation at 512 feet
t50 = 404 seconds
Time to reach50 consolidation
A-51
Figure A35 Piezo-dissipation at Sandpoint Bridge Crossing Idaho for depth z = 512 feet illustrating the
evaluation of characteristic time t50
It is also common to plot normalized excess porewater pressures relative to the measured initial Du2
that is obtained during penetration at the constant rate of 20 mms ie DuDui versus time Figure A36
shows a set of piezo-dissipation records for a bridge and embankment site in southern Louisiana
reported by the LTRC While the porewater pressure axis is arithmetic the time scale is often plotted on
either logarithmic or square root scales In either case the t50 is then found from the measured
dissipation curve when DuDui = 050
A-52
Figure A36 Set of monotonic piezo-dissipations taken at various depths for Courtableau Bridge
Site on Louisana State Highway 103 (Abu-Farsakh et al 2011)
Dilatory Dissipation Curves
In a number of situations a monotonic type dissipation does not occur on the onset but instead a dilatory
type curve is observed Here after stopping the push of the penetrometer the recorded porewater
pressures initially begin to rise and eventually reach a peak value then afterwards follow a phase where
the Du values decrease to hydrostatic (Figure A37) A full solution is available for both cases (Burns amp
Mayne 2002) Dilatory responses are most often associated with overconsolidated clays and silts
although occasionally are observed in fine-grained soils with low OCRs
The selection of the characteristic t50 value may found by using a square root of time plot to find the initial
u2 value by projection of the post-peak dissipatory data back to the ordinate axis as shown by the example
in Figure A38 In this example the initial u2 = 480 kPa and then the same procedures in Figure A35may
apply
A-53
Figure A37 Monotonic and dilatory porewater dissipation responses in soils
(after Burns amp Mayne 1998)
Figure A38 Method for obtaining t50 from dilatory dissipation curves
(Schneider amp Hotstream 2010)
A-54
Interpretation of Dissipation Test Data
There are a number of different solutions available for the interpretation of the coefficient of
consolidation (cv) from the piezocone dissipation curves For monotonic type response the strain path
method (SPM) developed at Oxford University (Teh amp Houlsby 1991) is well-recognized while the Georgia
Tech solution by Burns amp Mayne (2002) is based on spherical cavity expansion and critical state soil
mechanics (SCE-CSSM) and handles both monotonic and dilatory curves For both solutions a simplified
procedure can be recommended that relies on the aforementioned t50 value obtained from the measured
field data as given in Table A3 The units for cv are in cm2s as shown or equivalent
Table A3 Recommended procedures for calculating cv from t50 obtained in dissipation tests
Method Equation for coefficient of
consolidation cv
RemarksNotes Eqn
No
SPM
(Teh amp
Houlsby
1991) 50
2)(2450
t
Iac
Rc
v
t50 = measured time to reach 50
dissipation
IR = Gsu = undrained rigidity index
G = shear modulus
su = undrained shear strength
ac = penetrometer radius (ac = 178
cm for 10-cm2 cone ac = 220 for a
15-cm2 size)
A32
SCE-CSSM
(Burns amp
Mayne 2002) 50
7502 )()(0300
t
Iac Rc
v
A33
A-55
Evaluating rigidity index
Both the SPM and SCE-CSSM solutions require an estimate of the in-situ rigidity index (IR = Gsu) of the
soil If the results of SCPTU are available Krage et al (2014) have derived an expression for IR that
depends upon the small-strain shear modulus net cone tip resistance and effective overburden stress
250750
max50
8111
vonet
Rq
GI
s A38
where consistent units are input for Gmax qnet and svo The value of Gmax is obtained from B29 using the
measured shear wave velocity (Vs) and unit weight evaluated from B7
If the shear wave velocity is not measured it can be estimated from the CPT data A number of
methods have been reviewed for CALTRANS by Wair et al (2012) including one that is applicable
to sands silts and clays of low sensitivity that are inorganic and uncemented
119881 (119898119904) = [101 middot 119897119900119892(119902119905) minus 114]167 middot (100 middot 119891119904119902119905)03 A39 119904
where qt is in units of kPa (Hegazy amp Mayne 1995) An alternative relationship is given by Robertson amp
Cabal (2015)
A114 Hydraulic Conductivity
The hydraulic conductivity is a parameter that expresses the flow characteristics of the soil In
geotechnics it is also called the coefficient of permeability (k) and has units of cms or feetday
Through consolidation theory the hydraulic conductivity relates directly to the coefficient of
consolidation
A-56
A40 D
ck wv
where w = unit weight of water and D = constrained modulus
Therefore one approach to evaluating k can be from the site-specific cv obtained from the
dissipation tests using an estimate of D from one of the relationships given in Table A2
An approach developed for soft normally-consolidated soils that uses t50 directly to assess the magnitude
of k is presented in
Figure A39 (Parez and Fauriel 1988) The mean trendline through the wedges of soil type can be
approximated by
251
50 (sec)251
1)(
tscmkh
A41
A-57
Figure A39 Hydraulic conductivity versus dissipation time for 50 consolidation
(Parez amp Fauriel 1988)
A-58
APPENDIX B
List of Figures
Figure B1 Conventional methods vs direct CPT approach to shallow foundation response B-3
Figure B2 Direct bearing capacity relationship for embedded square and strip footings situated on clean
sands (after Schmertmann 1978) B-6
Figure B3 Direct CPT bearing factor Rk as function of footing width B and embedment depth from finite
element analyses (Tand et al 1995) B-6
Figure B4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio and sand
consistency (Eslaamizaad amp Robertson 1996) B-7
Figure B5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp Salgado (2005) in
terms of footing size relative density and base settlement B-8
Figure B6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and normalized
cone tip resistance (after Eslami amp Gholami 2005) B-8
Figure B7 Direct relationship between ultimate bearing stress in clays and measured cone tip resistance
(Schmertmann 1978) B-10
Figure B8 Direct relationship between ultimate foundation bearing stress in clays and net cone tip
resistance (after Tand et al 1986) B-11
Figure B9 Measured load-displacements for each of the five large footings at TAMU (Briaud amp Gibbens
1994) sand site B-12
Figure B10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU (Briaud amp
Gibbens 1994) footings B-13
Figure B11 Characteristic stress versus square root normalized displacements for TAMU footingsB-15
Figure B12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sandB-16
Figure B13 Summary of sand formation factors with corresponding CPT resistancesB-19
Figure B14 Normalized foundation stresses vs square root of normalized displacement for spread
footings on sand (modified after Mayne et al 2012) B-20
Figure B15 Definitions of capacity from three types of observed foundation load test responses
(modified after Kulhawy 2004) B-21
Figure B16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footingsB-22
B-1
Figure B17 Measured load vs displacement curves for 3 shallow foundations on clay at Baytown Texas
(Stuedlein amp Holtz 2010)B-25
Figure B18 Characteristic stress vs square root of normalized displacement for all three foundations at
Baytown site (Stuedlein amp Holtz 2010) B-26
Figure B19 Normalized foundation stress vs square root of normalized displacements B-27
Figure B20 Applied foundation stress vs CPT-calculated stress for 67 footingsB-28
Figure B21 Applied footing stress vs CPT calculated stress on logarithmic scales B-29
Figure B22 Summary graph and equations of direct CPT method for evaluating the vertical B-30
Figure B23 Foundation soil formation parameter hs versus CPT material index Ic B-32
Figure B24 Bearing capacity ratio qmaxqnet versus CPT material index IcB-33
Figure B25 Influence factor for rectangular foundations from elastic theory solution (Mayne amp
Dasenbrock 2017) B-34
List of Tables
Table B1 Direct CPT methods for bearing capacity of footings on clean sandsB-4
Table B2 Direct CPT methods for bearing capacity of footings on clays B-9
Table B3 Summary of large footings soil conditions and reference sources of database B-17
Table B4 Sources for shallow foundation performace B-34
B-2
B1 DIRECT CPT METHOD FOR SHALLOW FOUNDATIONS
B11 Introduction
The analysis of shallow foundations on soils is classically handled as a two-step and two-part set of
calculations involving (a) bearing capacity commonly reliant on limit plasticity solutions and (b)
settlement or more appropriately termed displacements that are assessed via elastic continuum theory
The utilization of cone penetration tests (CPT) provides the necessary geotechnical data for the site-
specific information on the subsurface conditions at the project site including the geostratigraphy and
evaluation of input parameters This is still a viable approach where the CPT readings are interpreted to
give the soil unit weight (γt) effective friction angle (ϕ) undrained shear strength (su) preconsolidation
stress (σp) and elastic moduli (D and E) for analysis
An alternative approach is the use of CPT results to provide direct assessments of bearing capacity
andor settlements A comparison of these two distinctly different and alternate paths is depicted in
Figure B1
Figure B1 Conventional methods versus direct CPT approach to shallow foundation response
A number of available direct CPT approaches for shallow footings are reviewed within this report
Moreover as the entire load-displacement-capacity response of shallow foundations to loading occurs
as a continuous nonlinear phenomenon a single direct CPT solution is presented for the behavior of
B-3
footings on sands (drained) and clays (undrained) The methods discussed herein are specifically to
address the general case of vertical loading of shallow foundations Additional more complex situations
that consider load eccentricity moments inclined forces sloping ground and other facets may be
handled using well-established procedures that are discussed elsewhere (eg Vesić 1975 Kulhawy et
al 1983)
This report focuses on foundations on granular soils andor soils exhibiting drained behavior since
Federal Highway Administration (FHWA) studies have concluded that less than 1 of shallow
foundations for highway bridges are placed on clay soils (Paikowsky et al 2010) One reason for this is
because soft-firm intact clays require a more complicated process that assesses both a short-term
analysis (undrained) as well as long-term analysis (drained) thus requiring undrained bearing capacity
undrained distortion displacements drained bearing capacity and drained primary consolidation
settlements
B12 Direct CPT Methods for Bearing Capacity of Footings on Sands
Classical bearing capacity solutions are based most often in limit plasticity theory albeit can be
formulated from limit equilibrium cavity expansion and numerical methods Because the limit
plasticity solutions are prevalent and adopt total stress analysis they are usually developed as either
fully drained (Δu = 0) or fully undrained (ΔVV = 0) where Δu equals excess porewater pressure and
ΔVV equals volumetric strain Paikowski et al (2010) provides an extensive review of various solutions
There are a number of direct CPT methods for the evaluation of the bearing capacity of footings and
shallow foundations Table B1 lists a variety of these direct CPT approaches for footings on sands In
the direct approach a one-step process is used to scale the measured cone penetrometer readings (ie
measured cone tip resistance (qc) total cone tip resistance (qt) sleeve friction (fs) andor measured
porewater pressure u2)) to obtain the ultimate bearing stress (qult) of the foundation
Table B1 Direct CPT methods for bearing capacity of footings on clean sands
Method Surface Footing Remarks and Embedment Notes
Meyerhof (1956) qult = qc (B12) middotcw qult = qc (B12)(1+DeB)cw
Note for silty sand reduce by 05
cw = 10 dry or moist sand cw = 05 submerged sand
Meyerhof (1974) qult = qc (B + D)40 with stresses in tsf
Presumably dry sands B = fdn width (feet) and D = depth (feet)
Schmertmann (1978)
NA (applies to embedded footings)
Square qult =055σatm(qcσatm)078
Strip qult =036σatm(qcσatm)078
See Figure A2
Embedment applies De gt 05(1+B) for Blt1m De gt 12 m for B gt1m
B-4
Canadian qult = Rk0middotqc Applied to FS = 3 Geotech Society where Rk0 = 03 where FS = factor of (CFEM 199 2) safety
Tand et al (1995) NA qult = Rkmiddotqc + σvo See Figure A3 where Rk = fctn(De B)
Frank and Magnan (1995)
qult = Rk0middotqc where Rk0 = fctn(sand
consistency) Loose Rk0 = 014
Medium Rk0 = 011 Dense Rk0 = 008
qult = Rk1middotqc + σvo where Rk1 = function (De B L amp
sand consistency) Factor Rk1 = Rk0[1+Rk206+04(BL) (DeB)]
and Rk2 = 035 (loose) 050 (medium) and 085 (dense)
Loose sand qc lt 5 MPa
Medium sand 8 MPa lt qc lt 15MPa
Dense sand qc gt 20 MPa
Eslaamizaad amp Robertson (1996)
qult = KΦmiddotqc See Figure A4
See surface footing equation KΦ = function (BDe shape and density)
Lee amp Salgado (2005)
qbL = βbcmiddotqc(AVG)
where qc averaged over distance B
Not addressed See Figure A5 for factor βbc = fctn(B DR
K0 and sB)
beneath the base
Eslami amp Gholami (2005
2006)
See embedded footing solution
qult = Rk1middotqc where Rk1 = function (ratio DeB
and normalized qcσvo) See Figure A6
Measured qc and qcσvo are geometric means over 2B deep
beneath footing
Robertson amp Cabal (2007)
qult = KΦmiddotqc with KΦ = 016
See surface footing equation
Briaud (2007) qult = KΦmiddotqc with KΦ = 023
Based on full-scale tests at Texas AampM
Mayne amp Illingworth
(2010) qult = 018 qc-Mean
Note qc -Mean is averaged CPT cone resistance over depth of
influence z = 15 B deep
Based on 30 footing load tests on 12
sands
Lehane (2013) qult = 016 qc-AVE Note qc-AVE is averaged CPT cone resistance within z (m) = [B (m)
]07
Based on 47 load tests
Table B1 Continued
Notes B = minimum footing width (or diameter) L = footing length De = embedment depth cw = water
table correction σvo = total overburden stress at bearing elevation and σvo = the effective vertical stress
B-5
Figure B2 Direct bearing capacity relationship for embedded square and strip footings
situated on clean sands (after Schmertmann 1978)
Figure B3 Direct CPT bearing factor Rk as function of footing width B and embedment depth
from finite element analyses (Tand et al 1995)
B-6
Figure B4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio
and sand consistency (Eslaamizaad amp Robertson 1996)
B-7
Figure B5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp
Salgado (2005) in terms of footing size relative density and base settlement
Figure B6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and
normalized cone tip resistance (after Eslami amp Gholami 2005)
B13 Direct CPT Methods for Bearing Capacity of Footings on Clays
For direct CPT evaluations of foundation bearing capacity in clays Table B2 summarizes some of the
well-documented methods that are available Generally these assume that the loading is fast enough
and the permeability of the clay is also sufficiently low such that an undrained (ie constant volume)
condition is maintained This is fine for short term loading of foundations particularly for soft to firm
intact clays However the undrained case is not permanent Given sufficient time excess porewater
pressures in the clay will eventually dissipate and hydrostatic conditions will return to equilibrium
During this dissipation phase primary consolidation will occur that results in drained settlements Also
a drained bearing capacity condition will prevail As the CPT is advanced at a constant rate of 20 mms
the recorded readings normally constitute undrained behavior from a direct measurement viewpoint
Thus direct CPT methods are normally focused at an undrained foundation response Nevertheless the
drained footing case can be addressed using CPT results via use of conventional analysis approach and
effective stress limit plasticity solutions that require piezocone penetrometers with porewater pressure
measurements
B-8
B
D)
L
B4060(3501320kc
Many of the earlier solutions were based on data from mechanical penetrometers andor older electric
cones that measure the uncorrected tip resistance (qc) It is now well-recognized that this measurement
must be updated to the corrected cone resistance (qt) because of porewater pressure effects that act on
the mechanics of the load cell and geometry of the particular penetrometer design The results are
particularly significant in clays silts and mixed soils that are soft to firm to stiff where excess porewater
pressures are recorded
Table B2 Direct CPT methods for bearing capacity of footings on clays
Method Footing Comments NotesRemarks
Meyerhof (1974) qult = αbcmiddotqc
where 025 le αbc le 050
Applicable to saturated insensitive clays under short-term loading
Cone tip resistance
measured by
mechanical CPT
Trofimenkov (1974)
Approximation
qult asymp (qc33)09
(Assuming FS = 3)
Strip footings on clays and sandy clays 06m le B le 15m 1m le zemb le 25m
Mechanical CPT
with qc and qult in
kgcm2
Schmertmann (1978)
qult = function (qc and foundation shape)
Relationship shown in Figure A7 for square and strip footings
Cone tip resistance
measured by
mechanical CPT
Tand et al (1986)
qult = Rkmiddot(qc - σvo) + σvo
Factor Rk obtained from Figure A8 accounts for surface and deep footings Separate Rk factor for intact and jointed clays
= (qc1middotqc2)05 qc
where qc1 is geometric mean from bearing elevation to 05B deeper and qc2 is geometric mean from 05B to 15B beneath foundation base
Mix of data from
mechanical CPT qc
and electric CPT qc
LCPC Method
(Frank amp Magnan 1995)
Footing a surface
qult = kc middotqc + σvo
where kc = 032
Embedment case
Note Methods above generally consider undrained bearing capacity
B-9
Figure B7 Direct relationship between ultimate bearing stress in clays and measured cone tip
resistance (Schmertmann 1978)
B-10
Figure B8 Direct relationship between ultimate foundation bearing stress in clays and net
cone tip resistance (after Tand et al 1986)
B14 Direct CPT Assessments of Settlement and Displacements of Shallow Foundations
Methods for evaluating the magnitude of foundation displacements (s) can be found using
elastic theory subgrade reaction methods spring models and numerical simulations The
classical approach to footing settlement calculations on sands is to utilize elastic theory
solutions (eg Poulos amp Davis 1974) which take on the form
119902∙119861∙119868∙(1minus1205842) 119904 = B1
119864119904
where q = applied footing stress and I = displacement influence factor from elasticity theory
The value of I depends upon foundation geometry footing rigidity embedment depth variation of soil
modulus with depth compressible layer thickness and other factors as discussed by Mayne amp Poulos
(1999) For instance for a flexible circular footing of diameter B resting on an infinitely deep soil
formation having a constant modulus Es with depth the influence factor I is equal to 1 for the center
point The results of in-situ field tests such as pressuremeter tests (PMT) standard penetration tests
(SPT) cone penetration tests (CPT) flat dilatometer tests (DMT) and other methods can be used to
ascertain the input value of soil modulus Es
Alternatively direct CPT methods have been investigated to provide a one-step assessment of
foundation displacements Selected methods for direct CPT evaluation of shallow foundation
settlements on granular soils is include DeBeers amp Martens (1957) Meyerhof (1965) DeBeer (1965)
Thomas (1968) Schmertmann (1970) Meyerhof (1974) Berardi amp Jamiolkowski amp Lancellotta (1991)
Robertson (1991) Lutenegger amp DeGroot (1995) Lehane amp Dougherty amp Schneider (2008) Gavin amp
Adekunte amp OrsquoKelly (2009) Mayne amp Illingworth (2010) and Uzielli amp Mayne (2012)
B141 Large Scale Footing Load Tests
The measured load-displacement response of shallow foundations is conducted in the field using either
stepped loading procedures or continuous rate of displacement methods Results from the well-
documented test program at the Texas AampM University (TAMU) site (Briaud amp Gibbens 1999 Briaud
2007) for five spread footings on sand are shown in Figure B9 Each footing clearly shows a nonlinear
B-11
behavior to loading over the range of testing This site is one of the national geotechnical
experimentation sites (NGES) funded by the National Science Foundation (NSF) FHWA and American
Society of Civil Engineering (ASCE)
Figure B9 Measured load-displacements for each of the five large footings at TAMU (Briaud
amp Gibbens 1994) sand site
B142 Generalized Direct CPT Method for Footing Response on Soils
The use of applied foundation stress versus normalized displacement curves (q vs sB) is generalized to
footings on sands silts intact clays and fissured clays A square root plotting of the normalized
displacements (sB) permits a single parameter characterization for each specific soil type that in turn
correlates with the net cone tip resistance The generalization is based on a statistical review of data
from large foundations (B gt 05m) involving 70 full-scale load tests with available CPT data For fine-
grained soils the data are from electronic piezocone tests where the proper correction from qc to qt has
been obtained to allow the best possible results The methodology permits an evaluation of the load-
displacement-capacity response of shallow square footings based on CPTs performed in the four types
B-12
of ground conditions For the TAMU footings the characteristic stress-normalized displacement
response is shown in Figure B10 where all five foundations can be represented by a single curve
Figure B10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU
(Briaud amp Gibbens 1994) footings
B143 Characteristic Stress-Displacement Curves
Fellenius amp Altaee (1994) recommended the concept of a characteristic stress (q) vs normalized
displacement (sB) curve for foundation response on a given soil formation later supported by Decourt
(1999) Briaud amp Gibbens (1999) Lutenegger amp Adams (2003) and Briaud (2007) In this approach the
measured load (Q) vs displacement from individual footings of various sizes that rest on the same soil
conditions collapse to a unified relationship given in Equation B2
119939119943 119956 119954119938119953119953119949119946119942119941 = 119938119943 (119913
) B2
B-13
where af and bf are empirical fitting coefficients (Decourt 1999 Uzielli amp Mayne 2011 2012)
In a review of measured load tests data from large spread footings situated on different sands it has
been suggested that Equation B2 can be reduced to a single parameter expression by use of square root
plotting where the exponent bf is equal to 05 (Mayne amp Illingworth 2010 Mayne et al 2012)
119956 119954119938119953119953119949119946119942119941 = 119955119956radic
119913 B3
where rs = fitted soil parameter from best fit line regression In the case of sands the coefficient rs
represents the supporting soil conditions including particle size relative density and fines content as
well as geologic origin aging and overconsolidation effects
The results from the five TAMU footings are presented in this format in Figure B11 that shows the
formation factor rs = 486 MPa for this sand deposit This single coefficient can be used to express the
nonlinear load versus settlement of any size footing on this sand site Moreover one criterion from
Eurocode for bearing capacity that is used by the European community identifies that stress (or load)
which corresponds to (sB) = 10 That is when the displacement equals ten percent of the footing
width then the capacity has been reached Therefore adopting this criterion for footings on sands
the ultimate bearing stress (qult) for footings on sand can be taken when (sB) equals 010 or when
square root of (sB) equals 0316 In that case this gives qult equal to 0316 middot rs For the TAMU site this
gives qult equal to 0316 times (486) which equals 153 MPa as illustrated by Figure B12
B-14
Figure B11 Characteristic stress versus square root normalized displacements for TAMU
footings
B-15
Figure B12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sand
B15 Footing Database
A database of spread footings and large plates was assembled which considers only full-size shallow
foundations (05 le B le 6 m) that rest on sands silts and clays Sites for the foundations were also subjected to CPT The inclusion solely of large footings are significant because results from small-scale
model footings exhibit scale effects In fact experimental centrifuge work and numerical simulation
studies have shown that bearing capacity factors are size-dependent especially for factor Nγ decreasing
with footing width B (Kimura et al 1985 Cerato amp Lutenegger 2007 Mase amp Hashiguchi 2009) A
direct approach based on full-size footings provides a more reliable means for foundation evaluation
The compiled database is given in Table B3 with a total of 70 large footings and plates These include a
listing of 34 foundations on 13 sands 11 footings on 4 silt deposits 13 footings or large plate load tests
on 6 intact clays and 12 foundations on 5 fissured clays Most of the footings were square or nearly
square (80) while the remaining were circular The largest foundation was a mat 5 m by 14 m in plan
loaded by stacking Kentledge concrete blocks to cause a bearing failure in the underlying soft clay For
sands the largest footing consisted of the 6-m square concrete pad In terms of equivalent square
footings mean footing sizes were 155 m (sands) 114 m (silts) and 126 m (clays)
B-16
Additional details on the individual load tests and site conditions are given elsewhere (Mayne 2009
Mayne amp Illingworth 2010 Uzielli amp Mayne 2011 2012 Mayne et al 2012 Mayne amp Woeller 2014)
Table B3 Summary of large footings soil conditions and reference sources of database
Sand Site Location Soil Conditions Footings Numbers
Shapes and Sizes
ReferencesSource
College Station
Texas Pleistocene sand 5 Square 1 15 25 33 m Briaud amp Gibbens (1999)
Kolbyttemon Sweden Glaciofluvial sand 4 Rect B = 06 12 17 24 m
Bergdahl et al (1985 1986)
Fittija Sweden Glaciofluvial sand 3 Rect B = 06 17 24 m Bergdahl et al (1984 1985)
Alvin West Texas Alluvial sand 2 Circular D = 235 m Tand et al (1994)
Alvin East Texas Alluvial sand 2 Circular D = 22 m Tand et al (1994)
Perth Australia Silceous dune sand 4 Square B = 05 and 10 m
Lehane (2008)
Grabo T1C Sweden Compacted sand fill
1 Square B = 046 m Long (1993)
Grabo T2C Sweden Compacted sand fill
1 Square B = 063 m Long (1993)
Grabo T3C Sweden Compacted sand fill
1 Square B = 080 m Long (1993)
Labenne France Dune sand 6 Square B = 07 and 10 m
Amar et al (1998)
Green Cove Florida Brown silty sand 1 Circular D = 182 m Anderson et al (2006)
Durbin South Africa
White fine sand 1 Square B = 609 m Kantley (1965)
Porto Portugal Residual clayey sands
1 Circular D = 12 m and 1 plate
Viana da Fonseca (2003)
Jossigny France Soft clayey silt 2 Square B = 1 m Amar et al (1998)
Tornhill Sweden Glacial Baltic till 3 Square B = 05 1 2 m Larsson (2001)
Vagverkel Sweden Stiff medium silt 3 Square B = 05 1 2 m Larsson (1997)
Vattahammar Sweden Brn-Gry layered silt 3 Square B = 05 1 2 m Larsson (1997)
B-17
Table B3 Continued
Clay Site Location Soil Conditions Footings Numbers Shapes and Sizes
ReferencesSource
Baytown Texas Fissured Beaumont 2 plates with d = 076 m Stuedlein amp Holtz (2008)
Belfast Ireland Soft clay silt ldquosleechrdquo
1 Square Pad B = 20 m Lehane (Geot Engr 2003)
Bothkennar Scotland Soft silty clay 2 Square Pads B = 22 24 m
Jardine et al (1995)
Bangkok Thailand Soft to stiff clay 4 Square 067 075 090 10 m
Brand et al (1972)
Haga Norway Stiff OC clay 2 Square Footings B = 10 m
Andersen amp Stenhammar (1982)
Rio Grande Brazil Sandy residual clay 3 Square 04 07 and 10 m
Consoli et al (1998)
Shellhaven England Soft estuarine clay Rectangular 5 m by 14 m Schnaid et al (1993)
Texas City A Texas Coast Fissured Beaumont 3 circular plates d = 058 m
Tand et al (1986)
Texas City B1 Texas Coast Fissured Beaumont 2 circular plates d = 058 m
Tand et al (1986)
Texas City B2 Texas Coast Fissured Beaumont 1 circular plate d = 058 m
Tand et al (1986)
Alvin Texas Texas Coast Fissured Beaumont 3 circular plates d = 058 m
Tand et al (1986)
For the series of footings on sands Figure B13 shows the summary of measure formation factors (rs) for
the 34 foundation load tests Also indicated on the graph are the corresponding mean qc values of the
sands averaged over 15middotB beneath the bearing elevations
Note also in sands that the qt and net resistance (qnet) are all very close to the measured qc since (a)
porewater pressure corrections for the a-net value are negligible in clean sands and (b) overburden
stresses are typically only 1 to 3 of the magnitude of qc This is especially true of shallow depths
Therefore for clean sands qnet is approximately qt which is approximately qc
B-18
Figure B13 Summary of sand formation factors with corresponding CPT resistances
As suggested by Briaud (2007) various in-situ measurements can be used to normalize the results from
cone penetration tests in this case Herein the qc from the CPT soundings in the sands are used to
normalize the footing stress axis as presented in Figure B14 It is evident that the results of the load-
displacement response of all 34 footings on 13 different sands can be captured by the characteristic
stress versus normalized displacement with the inclusion of the site-specific cone tip resistance In this
case the mean trend for all footings and sand sites indicates a simple expression shown in Equation B4
119954119943119952119952119957119946119951119944 119956 Clean quartz-silica sands = 120782 120787120790120787radic B4
119954119940 119913
B-19
Figure B14 Normalized foundation stresses vs square root of normalized displacement for
spread footings on sand (modified after Mayne et al 2012)
The results of measured force-displacement curves (Q vs s) from footing load tests are nonlinear and
thus the definition of capacity must be addressed Kulhawy (2004) identified three main curve types
that occur during load testing depicted in Figure B15 The most common response (Type A) shows no
clear peak and is observed in sands and silts as well as slow loading of insensitive clays and fissured
clays In this case ldquocapacityrdquo can be defined by the Eurocode corresponding to the load (or stress) when
sB=10 (Amar et al 1998) For foundations situated on many clays a plateau capacity is reached as
depicted as Type B response In rare cases where sensitive clays or structured soils are encountered a
Type C relationship occurs whereby the load-displacement curve reaches a peak followed by strain
softening In the one case observed in this study for Type C loading (Haga clay) the corresponding
ldquopeak capacityrdquo was reached at a pseudo-strain of sB = 4 as shown in Figure B16 followed by some
strain softening In the cases of soft clays at Belfast and Bothkennar a plunging type (plateau failure)
was achieved at around sB=7
B-20
Figure B15 Definitions of capacity from three types of observed foundation load test
responses (modified after Kulhawy 2004)
B-21
Figure B16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footings
In the case of fine-grained soils that develop excess porewater pressure during cone penetration the
use of the net total cone resistance (qtnet) which equals the cone tip resistance minus the total stress
must be considered (Lunne et al 1997) As noted earlier the correction of CPT data in sands is minimal
because of the low value of penetration porewater pressures and the fact that total overburden stress is
small relative to the cone resistance especially for shallow soundings Thus the use of qtnet may be
substituted for qc into the trend of Figure B1
B16 Undrained and Drained Capacity
The footings on sands typically show drained response with no excess porewater pressures developed
during loading For the foundations situated on silts analyses by Larsson (1997 2001) concluded that
these too show drained conditions For shallow foundations on clays the entire range of drainage cases
are possible ranging from undrained to partially-drained to fully-drained depending upon the applied
rate of loading relative to the permeability of the geomaterial If sufficient data were available for each
of the clay sites then the recent evaluation of drainage condition could be assessed using normalized
velocity criteria (Chung et al 2006) However this was not possible with the current data set Instead
B-22
the test loadings of footings on clays have essentially been assumed to primarily occur under undrained
conditions following recommendations by Tand et al (1986) For this case a limiting bearing stress can
be calculated from bearing capacity analyses and related directly to the CPT readings
From classical bearing capacity calculations using limit plasticity solutions the ultimate foundation stress
is shown in Equation B5
= 119925119940 ∙ 119956119958 B5 119954119958119949119957
where Nc equals the bearing factor for constant volume (514 for strip and 614 for square or circular
plan) and su equals the undrained shear strength of the clay For the case of soft to firm clays of low to
medium sensitivity the operational strength can be related to the preconsolidation stress (Mesri 1975
Jamiolkowski et al 1985) as shown in Equation B6
prime 119956119958 = 120782 120784120784120648119953 B6
Finally the effective preconsolidation stress has been linked directly to the net cone resistance (Chen amp
Mayne 1996 Demers amp Leroueil 2002) as shown in Equation B7
prime 120648119953 = 120782 120785120785(119954119957 minus 120648119959119952) B7
Combining Equations (B5) (B6) and (B7) results in direct expressions for undrained bearing capacity on
intact clays from CPT results as shown in Equations B8 and B8-2
Strip footing 119954119958119949119957 = 120782 120785120789120785(119954119957 minus 120648119959119952) B8
B-23
Squarecircle 119902119906119897119905 = 0445(119902119905 minus 120590119907119900) B8-2
Equation (B8-2) is in excellent agreement with Tand et al (1986) database study for the case of intact
clays and footings with no embedment Their study also provided a lower relationship for foundations
situated on jointed clays specifically recommending that the bearing stress not exceed 030 qtnet
Review of the available data in Figure 3 shows this to be rather conservative and a more realistic value
may be taken at 040 qtnet for fissured clays
B161 Normalized Undrained Footing Response
The characteristic stress vs normalized displacement curves for footings on clays exhibiting undrained
conditions can be verified using a relatively recent case study on Beaumont clay by Stuedlein amp Holtz
(2010) The Baytown Texas site was investigated using an exploration program of soil borings
piezocone soundings and laboratory testing The field testing included a large square footing (B = 276
m) and two small circular plates (D = 076 m) The measured load-displacement responses for all three
foundations are presented in Figure B17 Of course the large footing carried considerably higher loads
than the two small plates on the same soil deposit Yet using the aforementioned characteristic stress
(q) vs square root of normalized displacement (sB) essentially gave a unique relationship for all 3
foundations as presented in Figure B18 In this case utilizing the form of Equation B2 the
characteristic coefficient rs is 177 MPa for this deltaic clay formation
B-24
Figure B17 Measured load vs displacement curves for 3 shallow foundations on clay at
Baytown Texas (Stuedlein amp Holtz 2010)
B-25
Figure B18 Characteristic stress vs square root of normalized displacement for all three
foundations at Baytown site (Stuedlein amp Holtz 2010)
B17 General Direct CPT Method
In a manner analogous to the normalization of applied foundation stresses shown in Figure B14 for
footings on sands a generalized direct CPT method for shallow foundations on different soils can be
made in Equation B9
120782120787 119956 119954 = 119945119956 ∙ 119954119957119951119942119957 ∙ (119913
) lt 119954119940119938119953119938119940119946119957119962 B9
where qcapacity is equal to the foundation bearing capacity of the ground and hs is equal to the empirical
fitting term that depends on soil type Specifically hs equals 058 for sands 112 for silts 147 for
fissured fine-grained soils and 270 for clays A summary of the normalized stress versus the normalized
displacement curves for these four soil types is presented in Figure B19 where the data fitting to obtain
the hs parameter on clays are limited to (sB) lt 004 corresponding to undrained loading The data on
B-26
silts and sands are considered fully drained whereas the fissured clay subset may be partially drained to
undrained The statistical measures for obtaining the fitted parameter hs are quite good as shown in
Figure B20 with the coefficient of determinations (r2) values of 0947 for sands 088 for silts 0935 for
fissured and 0925 for intact clays Additional statistical evaluations on the database are provided by
Uzielli amp Mayne (2011)
Figure B19 Normalized foundation stress vs square root of normalized displacements
B-27
Figure B20 Applied foundation stress vs CPT-calculated stress for 67 footings
The same data are shown on Figure B21 in a log-log plot to emphasize the early parts of the load-
displacement responses of these footings The best fit line from regression analyses shows excellent
statistical correlations As detailed earlier the undrained bearing capacity of square or circular footings
on clays can be taken as approximately 040 times qtnet For silts and sands that experience drained
loading with no excess porewater pressures being developed and no clear peak value for capacity the
(sB) =10 criterion can be used In this case Equation 7 can be used directly to obtain stress q equal to
qcapacity when (sB)05 is 0316
An alternate approach is presented in Figure B22 summarizing the direct CPT method for estimating the
load-displacement-capacity of shallow square footings on four soil types The maximum values of soil
bearing capacity (qmax) can be capped at stresses corresponding to a percentage of the average
measured qtnet determined over the depth interval from the foundation bearing elevation to 15middotB
below the foundation In such an approach the foundation bearing capacity can be taken as
(qcapacityqtnet) of 02 for sands 035 for silts 040 for fissured and 045 for intact clays Using the assigned
values of the hs parameter the corresponding cut-off for the general stress-normalized displacement
relationship given by Equation 7 can be established specifically giving corresponding values of the
B-28
normalized displacements which can also be considered as pseudo-strains equal to (sB) max of 4 for
clays 7 for fissured clays 10 for silts and 12 for sands
Figure B21 Applied footing stress vs CPT calculated stress on logarithmic scales
B171 CPT Soil Material Index Ic
The soil behavioral type can be assessed indirectly by CPT using a material index (Ic) as defined by
Robertson (2009a b) shown in Equation B10
119920119940 = radic[120785 120786120789 minus 119949119952119944119928119957119951]120784 + [120783 120784120784 + 119949119952119944119917119955]120784 B10
B-29
Figure B22 Summary graph and equations of direct CPT method for evaluating the vertical
where Qtn equals the stress-normalized cone tip resistance and Fr equals the normalized sleeve friction
calculated from the cone penetrometer readings as shown in Equations B11 and B12 respectably
(119954119957minus120648119959119952)120648119938119957119950 = 119951 B11 119928119957119951 prime (120648119959119952120648119938119957119950)
119943119956 119917119955() = 120783120782120782 ∙ B12 (119954119957minus120648119959119952)
where σatm equals a reference stress equal to atmospheric pressure (σatm = 1 atm asymp 1 bar asymp 100 kPa) In
the initial evaluation the exponent n is set to n = 1 to find the soil behavioral type (SBT) based on a 9-
zonal chart as shown in Figure 23 The value of exponent n varies with soil type ranging from n = 1 in
intact clays and decreasing with increasing grain size to around approximately 075 in silts and
B-30
approximately 05 plusmn 02 in clean sands The appropriate value of exponent n is found by iteration until
conversion using the relationship (Robertson 2009b) shown in Equation B13
prime 120648119959119952 119951 = 120782 120785120790120783 ∙ 119920119940 + 120782 120782120787 ( ) minus 120782 120783120787 le 120783 120782 B13 120648119938119957119950
As the distinction that footings on intact clays subjected to fast loading are behaving primarily under
undrained conditions (ie constant volume) Robertson (2009a) suggested that this occurs when Ic gt 27
while in contrast Ic lt 25 corresponds more or less to drained loading (ie no excess porewater
pressures)
The CPT material index can be used to identify soil type and the corresponding characteristic hs value for
estimating footing load-displacement response In a number of the case studies investigated the results
of the sleeve friction readings were also available to allow the calculation of Ic with depth at these sites
Figure B23 shows the tentative relationship between the values of the hs parameter plotted versus the
corresponding CPT material index with an approximate trend given by Equation B14
120784120785 119945119956 = 120784 120790 minus 120783120787 B14
119920119940 120783+( ) 120784120786
B-31
Figure B23 Foundation soil formation parameter hs versus CPT material index Ic
Soil behavioral type from the Ic ranges given by Robertson (2009b) and are shown in Figure B23 with an
overall general agreement with the known soil classifications In summary the CPT can provide an
evaluation of the load-displacement-capacity response directly via Equation B15 which may be of
interest in mixed soil types such as silty sands sandy silts and the like
120784120785 Footing stress 119954 = 119954119951119942119957 ∙ radic(119956frasl119913) ∙ [120784 120790 minus
)120783120787] B15
120783+(119920119940frasl120784120786
As discussed earlier foundation bearing capacity may be taken by a limiting value of pseudo-strain (sB)
max or by qmax as specified as a percentage of qtnet This definition of bearing capacity can be seen in
comparison to soil type as seen in Figure B24
B-32
Figure B24 Bearing capacity ratio qmaxqnet versus CPT material index Ic
B172 Rectangular Foundations
This same elastic solution from Giroud (1968) for a CPT method for square and circular foundations were
utilized for rectangular foundations The study covered rectangular distortions (AB) ranging from 1
(square) to very long foundations with AB equal to 20 where A represents the foundation length and B
represents the foundation width (Mayne amp Dasenbrock 2017) The influence factor for rectangular
shaped foundations is given in Figure B25 and the expression in Equation B16
= (119912119913)120782120785120786120787 119920119912119913 B16
B-33
Figure B25 Influence factor for rectangular foundations from elastic theory solution (Mayne
amp Dasenbrock 2017)
Including 32 full scale load tests on sands results from an additional 98 shallow foundations have been
reported in previous studies The foundations were primarily square or circular and among these include
large spread footings with measured settlements and bearing capacity Table B4 shows a summary of
the number of footings and footing sizes used in the study The range of length to width ratios varied
from 1 lt (AB) lt 23 with an average value of (AB) equal to 238 The embedment depth to footing
width ranged from 0 lt DeB lt 222 and averaged 046
Table B4 Sources for shallow foundation performace
Source of Data No of
Footings
Mean B
(m)
Max B
(m)
Min B
(m)
Mean A
(m)
Max A
(m)
Min A
(m)
Mayne et al (2012) 32 149 609 046 149 609 046
Lehane (2011) 8 041 027 060 041 060 027
Gifford et al (1987) 17 846 1588 527 1591 3596 701
Jeyapalan amp Boehm
(1986)
13 1022 2892 400 1671 3049 640
B-34
Schmertmann
(1970)
31 945 5608 061 1288 8668 061
Papadopoulos
(1992)
29 870 3600 100 1300 7290 100
Total 130 671 5608 027 1012 8668 027
The solution for shallow foundation settlements given back in Equation B1 combined with
the expression in Equation B16 takes on the form
119954∙119913∙119920119912119913∙(120783minus120642120784) 119956 = B17
119916119956
Combining the direct CPT equation developed from the 32 full scale load tests (Equation B15)
together with the influence factor for rectangular shaped foundations becomes
minus120782120785120786120787 120784120785 119912 Footing stress 119954 = 119954119951119942119957 ∙ radic(119956frasl119913) ∙ [120784 120790 minus
)120783120787] ∙ [ ] B18
120783+(119920119940frasl120784120786 119913
B18 Conclusions
A direct CPT method for square rectangular and circular shallow footings is developed using a database
of 166 full-scale field load tests where the minimum footing size of an equivalent square width of 05 m
is established to avoid scaling issues The use of load vs displacement is generalized by characteristic
stress vs normalized displacement (sB) and further simplified by a square root plotting technique that
captures the ground response in a single parameter that is related to the qtnet Data are grouped
according to four main soil categories sands silts fissured clays and intact clays It is generally believed
that the footings on sands and silts exhibit fully drained behavior while in intact clays an undrained
response occurs under conditions of constant volume
B-35
The summary equation for evaluating the vertical stress-displacement-capacity of square rectangular
and circular footings is given by
minus120782120785120786120787 119956 119912 119954119943119952119952119957119946119951119944 = 119945119956 ∙ 119954119957119951119942119957 ∙ radic ∙ ( ) lt 119954119950119938119961 B19
119913 119913
where the parameter hs also relates directly to Ic The foundation capacity depends upon the mode of
failure as well as drainage conditions and can be taken as a fraction of qtnet where qmaxqtnet is 020 in
sands 035 in silts 040 in fissured clays and 045 in intact clays as shown in Figure 24 Alternatively the
capacity can be defined using a limiting pseudo-strain given by (sB) max of 4 in clays 7 in fissured
10 in silts and 12 in sands
B-36
APPENDIX C
List of Figures
Figure C1 Components of Axial Pile Capacity C-3
Figure C2 Selected alpha relationships as function of the undrained shear strengthC-5
Figure C3 Pile side friction coefficient β in terms of effective friction angle and overconsolidation ratio
C-11
Figure C4 Concept of Direct versus Rational Method for CPT evaluation of axial pile capacityC-14
Figure C5 Soil behavior type using CPT via original UniConeC-19
Figure C6 Soil behavior type using CPT via Modified UniCone C-19
Figure C7 Nine-zone soil behavioral soil type using normalized piezocone parameters C-21
Figure C8 Summary of Modified UniCone Method for direct CPT assessment of axial pile capacity C-22
Figure C9 Elastic continuum solution for axial pile response under compression and tension loadingC-24
List of Tables
Table C1 Alpha Methods for Axial Pile Side Friction where fp = αmiddotsu C-5
Table C2 Beta Methods for Axial Pile Side Friction where fp = βmiddotσvoC-9
Table C3 Selection of Direct CPT Methods for Axial Pile CapacityC-15
C-1
C1 EVALUATING DEEP FOUNDATION RESPONSE FROM CONE
PENETRATION TESTS
C11 Introduction
The axial response of deep foundations includes the load-displacement-capacity and axial load transfer
when driven pilings and drilled shafts are loaded in compression and uplift The use of cone penetration
testing (CPT) especially seismic piezocone tests (SCPTu) are advantageous since they provide
information on the subsurface soils and their geomaterial properties The SCPTu provides at least four
measurements on soil behavior with depth including (a) cone tip resistance qt (b) sleeve friction fs (c)
penetration porewater pressure u2 and (d) shear wave velocity Vs This offers the opportunity to
determine the geostratigraphy unit weight effective overburden stress shear strength stress state
and stiffness of the ground from a single exploratory sounding thus values in economic expediency
and reliability The axial pile capacity can be calculated based on static equilibrium of forces acting along
the sides and base of the pile foundation The displacement of the pile can be ascertained using
elasticity theory from closed-form solutions boundary elements andor finite element analyses Load
transfer occurs along the length of the pile and only a portion of axial forces are transmitted to the base
or toe or tip of the pile These too can be assessed using elasticity solutions
An alternate approach to pile capacity involves the utilization of direct CPT methods available
since the 1970s but in the past 10+ years a number of new and statistically reliable algorithms
have been developed which can be implemented for highway design and construction
C111 Axial Pile Capacity
The axial compression capacity of a single pile foundation is composed of a shaft or side component and
end-bearing component at the base as depicted in Figure 1 For a circular pile the side capacity (Qs) is
determined from the unit side friction (fp) acting along the surface area of the shaft which is As = πmiddotdmiddotL
where d = pile diameter and L = length embedded below grade
If the magnitude of side friction is uniform and constant with depth the side capacity is simply
C-2
119928119956 = 119943119953 middot 119912119956 C1
Moreover however many piles extend through multiple layers and a summation of unit side
frictions acting on various pile segments must be tabulated over the length of the pile as
suggested by Figure C1
Figure C1 Components of Axial Pile Capacity
The unit-end bearing resistance (qb) acts over the base of the pile tip where the area of a circular pile is
given by Ab = πmiddotd24 For piles in compression loading the base capacity is determined from Equation
C2
119876119887 = 119902119887 middot 119860119887 C2
and for piles in tension (or uplift) it is normally taken that Qb = 0
C-3
C112 Pile Unit Side Friction
Several different approaches can be adopted for evaluating the unit pile side friction (fp) prior to full-
scale load testing and construction (Poulos amp Davis 1980 ONeill 2001) The most common types
including the alpha and beta methods as summarized in Tables 1 and 2 respectively In the alpha
method an empirical coefficient (α) is applied to the undrained shear strength (su) of clay soils to
obtain
119891119901 = 120572 middot 119904119906 C3
An illustrative example of alpha curves is shown in Figure C2 A difficulty with the alpha
method is the evolution of many variants and changes to the expressions for its estimation It is
also restricted in its use for specific types of piles in clays and fine-grained soils
C-4
Figure C2 Selected alpha relationships as function of the undrained shear strength
Table C1 Alpha Methods for Axial Pile Side Friction where fp = αmiddotsu
C-5
Reference
Source
Alpha Equation Remarks
Tomlinson
(1957)
2716801102
uu ssAll pile types (steel
concrete timber)
Note su in ksf
Tomlinson
(1957)
2616201102
uu ssConcrete piles
Note su in ksf
McClelland
(1974)
Kerisel α = 07 - 031middotln(su)
Peck α = 10 - 02middotsu
Woodward α = 091middot(su)091
Driven piles in clay
Note su in ksf
Semple
(1980) )1(511
)1(1
OCR
OCR Summary from 9 series of
pile load tests
American
Petroleum
Institute (API
1981)
For suσvo le 035 α = 10
For suσvo gt 080 α = 05
Otherwise α = 1 + 1111middot (035 - suσvo)
Driven steel pipe piles
Tomlinson
(1986)
1 α = 55 (su)-091
2 α = 785 (su)-102
3 α = 605 (su)-100
where 75 lt su (kPa) lt 210
1 Driven piles
2 Bored piles
3 Driven piles in till with L
lt 10middotd
API (1987) 1 α = 10 for su lt 25 kPa Driven piles in clay other
than Gulf of Mexico
C-6
2 α = 125 - 001middotsu for 25 lt su lt 75 kPa
3 α = 050 for su gt 75 kPa
Kulhawy amp
Jackson
(1989)
α = 021+026middot(σatm su) le 1 Analyses of 106 drilled
shaft foundations
Table C1 Continued
API (1989) 05 For suσ vo le 1 α =
su radic frasl prime σvo
minus025 su For suσ vo gt 1 α = 05 ∙ ( frasl prime ) σvo
Driven steel pipe piles
Kulhawy amp
Jackson
(1989)
OCR
OCR
sin50
tan)sin1( sin
Λ = (1-CsCc) asymp 080 for
insensitive clays
Nowacki et al
(1996)
minus05 su For suσvo le 07 α = 05 ∙ ( frasl prime ) σvo
minus02 su For suσvo gt 07 α = 055 ∙ ( frasl prime ) σvo
Driven pile foundations
Kolk and van
der Velde
(1996)
α =09middot FL middot (suσvo)-03 lt 10
FL = [ (L - z)d]-02 = length term
L = pile length
Note su obtained from
UU lab tests
Applicable to driven piles
in clays
C-7
Knappett amp α = 1 for su le 30 kPa Non-displacement piles in
Craig (2012) fine-grained soils α = 116 - (su185) for 30 kPa le su le 150 kPa
α = 035 for su gt 150 kPa
Knappett amp
Craig (2012) α = 055 ∙ 02 40
( ) (LfraslD)
∙ su ( frasl prime σvo
minus03
) Displacement piles in fine-
grained soils
z = depth at point considered
d = outside diameter of pile
Karlsrud et al
(2005 NGI
Method)
α = function (suσvo and plasticity index) Driven pilings
Fleming et al
(2009) 05 minus05 su su For su σvo le 1 120572 = ( frasl prime ) ∙ ( frasl prime )
σvo NC σvo
05 minus025 su su For su σvo gt 1 α = ( frasl prime ) ∙ ( frasl prime ) σvo NC σvo
For driven piles in clay
where (suσvo)NC is the
normalized shear strength
for normally-consolidated
clay
Brown et al
(2010)
α = 0 for 0 lt z le 5 feet
α = 055 for z gt 5 feet and (suσatm) le 15
α = 055-01middot(suσatm -15) for 15le (suσatm) le 25
Drilled Shafts and Bored
Piles
Table C1 Continued
C-8
OCRK )sin1(0
Karlsrud
(2012)
α = function (su σvo and plasticity index) Driven pilings
Note as reported by Karlsrud (2012)
In the beta method the coefficient β is applied to the effective overburden stress (σvo) at the point of
concern along the pile length
prime = 120573 C4 119891119901 middot 120590119907119900
Table C2 Beta Methods for Axial Pile Side Friction where fp = βmiddotσvo
Reference Source Beta Equation Remarks
Burland (1973) β = (1-sinϕ) middot tan ϕ Piles in NC clays
Meyerhof (1976) β = Kmiddottanδ where K = K0 = for NC clays
K = 15middotK0 for OC clays
Poulos amp Davis
(1980)
β = (1-sinϕ) middot tan ϕ middot OCR05 Piles in OC stiff clays
Table C2 Continued
C-9
Kulhawy amp Jackson
(1989)
β = (1-sinϕ) middot tan ϕ middot OCRsinϕ Piles in quartz sands and insensitive
clays
ONeill (2001) β = (1-sinϕ) middot tan θ middot OCRsinϕ Drilled shafts where θ = (δϕ)middotϕ and
δ = interface friction between pile and soil
Fleming et al
(2009)
β = Kmiddottanδ K = lateral stress coefficient and δ =
friction angle between soil and pile
material
Karlsrud (2012) β = fctn (OCR and PI) PI = plasticity index of the clay ()
Mayne and Niazi
(2017)
β = CM middot CK middot K0 middot tanϕ
where K0 = (1-sinϕ) middot OCRsinϕ for soils
that are virgin loaded then unloaded
CM = pile material factor = 1 (drilled
augered) 09 (prestressed or precast
concrete) 08 (timber) and 07
(steel) and CK = pile installation factor
= 09 (bored or augered) 10 (low
displacement eg H-pile or open-end
pipe) and 11 (driven high
displacement eg prestressed
concrete closed-end pipe)
As the beta approach applies to all types of soils (gravels sands silts and clays) and more or
less has remained unchanged since its advent circa 1970 it has been selected for further
discussion herein In the direct form for calculation of unit pile side friction the expression is
given by
119874119862119877119904119894119899120601 prime prime 119891119901 = 119862119872 ∙ 119862119870 ∙ (1 minus 119904119894119899120601prime) middot ∙ 119905119886119899120601prime ∙ 120590119907119900 C5
C-10
where CM is a soil-pile interface friction coefficient and CK = pile installation factor Values of CM are in
the range 07 lt CM lt 10 depending upon pile material (steel wood concrete) and ranges for CK vary
09 lt CK lt 11 depending upon method of installation (auger drill driven) as detailed in Table 2
The value of K0 is limited to the passive stress coefficient which for the simple Rankine case is given by
KP = (1+sin ϕ) (1 - sin ϕ ) Thus there is a maximum value of overconsolidation ratio for which
Equation C5 applies given by OCRlimit = [(1+sin ϕ ) (1 - sin ϕ )2] (1sinϕ)
The corresponding graph in Figure C3 shows the relationship for β in terms of ϕ and OCR
Figure C3 Pile side friction coefficient β in terms of effective friction angle and overconsolidation ratio
C113 Pile Unit End Bearing
C1131 Theoretical Considerations
The evaluation of the end bearing resistance of pile foundations is commonly determined using
limit plasticity theory where
prime Drained loading = C6 119902119887 119873119902 middot 120590119907119900
C-11
Undrained loading 119902119887 = 119873119888 middot 119904119906 C7
where the value of σvo is calculated at the depth z = L and the value of su is the average undrained shear
strength from z = L to a depth z = L + d beneath the pile tip For a circular pile the limit plasticity solution
for undrained loading gives (Vesic 1977) a value Nc = 933 while a deep strip foundation would employ a
value of Nc = 824 For drained loading the expression for Nq for a deep foundation is given by (Vesic
1977)
)]arctan()sin1(tan21[)](tan1[sin1
sin1)tanexp( 2 BLABNq
C8
where A and B are the pile plan dimensions (for a circular pile A = B) and L = pile length For a
circular pile this can be approximated by
asymp 077(120601prime75deg) 119873119902 C9
over a range of effective friction angles 20deg le ϕ le 45deg
C1132 Practical Considerations
For drained loading of sands the full calculated end bearing capacity will never be realized because it
would require the pile to move a distance equal to its diameter This is beyond practical use and
therefore the limit plasticity solutions must be clipped to a fraction of the calculated value Another
reason for using a reduced end-bearing capacity is due to strain incompatibility since the side capacity is
mobilized early while the end-bearing is engaged much later So to have compatible values of Qs and
Qb the qb must be reduced using the following guidelines (Randolph 2003 Mayne 2007)
prime prime C10 119902119887 = 119873119902 ∙ 120590119907119900 ∙ 119891119909
where fx = strain incompatibility factor
fx = 010 for drilled shafts augered cast-in-place and bored piles
fx = 020 for driven low displacement piles (opened-ended steel pipe and H-piles)
fx = 030 for driven high-displacement piles (ie solid piles PSC and closed-ended pipe)
C-12
For undrained loading of circular piles in compression no reduction of end-bearing is necessary
therefore
119902119887 = 933 middot 119904119906 C11
C12 Direct CPT Methods for Pile Capacity
In the direct CPT method the penetrometer readings are scaled directly via specified algorithms to
obtain the pile unit side friction and end-bearing as depicted in Figure C4 As many as 40 different
direct CPT methods have been developed over the past five decades as summarized by Niazi amp Mayne
(2013) Starting circa 1970 many of these early methods relied on hand-recorded data from field
mechanical CPTs where only qc data were obtained at 20 cm intervals or later with mechanical readings
of both qc and fs using special sets of inner and outer rods that recorded vertical load for tip and loads
for tip plus sleeve in alternating increments Also early pile load test data were often obtained from
top-down measurements of load-displacement
C-13
Undrained qb = Nc su
Qside = S (fp DAs)
QTotal = Qs + Qb - Wp
fp = cmck Ko svorsquo tanrsquo
Qbase = qb Ab
Drained qb = Nq svorsquo
Method Two Rational or ldquoIndirectrdquo Method
Method OneldquoDirectrdquo CPT Method
(Scaled Pile)
fp = fctn (soil type pile
type qt fs and u2)
qb = fctn (soil type qt-u2)
qb = unit end bearing
unit sidefriction fp
OCR su Ko t DR rsquo
AXIAL PILE CAPACITY FROM CPT READINGS
Figure C4 Concept of Direct versus Rational Method for CPT evaluation of axial pile capacity
Beginning in the mid-1990s as the modern electric piezocone (CPTu) was implemented the newer
equipment offered improved data and better resolution because of the additional reading of porewater
pressures and correction of raw measured qc to total resistance qt In addition the use of electronic
digital data collection and field computers proved superior in field measurements and recordings As a
consequence several reliable CPT methods for axial pile capacity have been developed Moreover
parallel improvements in full-scale pile load testing occurred and now include modern strain gage
instrumentation digital data recording and automated testing procedures Also the testing can be
single direction (compression or tension) or bi-directional as in the Osterberg cell
C-14
Table C3 provides a selection of recent direct CPT methods that have become available over the
past two decades A number of these (namely ICP NGI UWA and Fugro) were funded by the
offshore industry because of the growth of oil amp gas reserves and windfarm installations thus
necessitating increased concerns on risk probability and reliability in the site investigations for
offshore platforms and design of large driven monopile foundations These direct CPT methods
are often statistically based on large datasets compiled from full-scale load tests made
worldwide (eg Schneider et al 2008)
Table C3 Selection of Direct CPT Methods for Axial Pile Capacity
Method Pile Types Soil Types References CPT
data
Additional
parameters
needed
Unicone Method all types Sands silts
clays
Eslami and
Fellenius
(1997)
Fellenius
(2009)
qt fs u2
KTRI = Kajima
Technical
Research
Institute
all types Sands
mixed clays
Takesue et al
(1998)
fs and u2 No guidance given
on end bearing
resistance
Imperial College
Procedure (ICP)
OE and CE Sands Chow et al
(1997 PhD)
Jardine et al
(2005)
qt Interface friction
angle (δ) from
ring shear tests
Correlated to
mean grain size
(D50)
C-15
OE and CE Clays Jardine et al
(2005)
qt Interface friction
angle (δ)
correlated to PI
Table C3 Continued
NGI Method
(Norwegian
Geotechnical
Institute)
OE and CE Sands Clausen et al
(2005)
qt DR from qt
Clays a Alpha
method
(Karlsrud et al
2005 2012)
b Beta
method
(Karlsrud
2012)
qt for su
qt for
OCR
PI = plasticity
index ()
PI = plasticity
index ()
Fugro Method OE and CE Sands Kolk et al
(2005)
qt
Clays Van Dijk and
Kolk (2011)
qt
OE and CE Sands Lehane et al
(2005)
qt IFR = infilling ratio
related to amount
of plugging
C-16
Purdue LFRD Drilled
shafts
Sands Basu amp
Salgado
(2012)
qt LRFD = load
resistance
factored design
Enhanced Various Sands silts Niazi and qt u2 Based on 330 load
Unicone Method clays and Mayne (2015 and fs tests including
mixed soils 2016) bored augered
jacked and driven
piles
UWA (Univ of
Western
Ra = pile
roughness
Australia) Clays Lehane et al
(2012)
qt Interface friction
angle (δ) from
ring shear tests
and also method
without δ
HKU Method
(Hong Kong
Univ)
OE and CE
(end
resistance
only)
Sands Yu and Yang
(2012)
qt End bearing only
PLR = plug length
ratio
Note PLR can be
estimated from
OE inner diam
Table C3 Continued
Note previously called Marine Technical Directorate (MTD)
C-17
Many of the offshore CPT methods relate to driven piles either in sand or clay as that is
common deep foundation for those purposes Of particular interest are the Unicone and
Modified Unicone Methods since they use all three readings of the piezocone (qt fs and u2)
and address a variety of pile foundation types
C13 Modified UniCone Method
The modified UniCone Method was developed based on a total 330 pile load tests which were
associated with SCPTu data during their site investigations (Niazi and Mayne 2015 2016) This
represents a threefold increase over the original UniCone database that was built upon data
from 106 pile load tests (Eslami amp Fellenius 1997)
For the original UniCone algorithms use is made of the effective cone resistance (qE)
119902119864 = 119902119905 minus 1199062 C12
and a chart of qE vs fs provided an approximate soil classification in five distinct groups as shown by
Figure C5 Later in the modified approach a better delineation of the larger dataset gave soil
subclassifications as indicated by Figure C6
C-18
Figure C5 Soil behavior type using CPT via original UniCone
Figure C6 Soil behavior type using CPT via Modified UniCone
C-19
In the modified approach the 9-zone normalized soil behavioral type (SBTn) is ascertained using CPT
data in conjunction with the Robertson (2009) charts as shown in Figure C7 As discussed in Appendix
A and B the SBTn method uses the normalized cone resistance (Qtn) normalized sleeve friction (Fr())
and CPT material index Ic This permits a much wider range in the calculated pile side friction because fp
is related as a continuous curve with Ic rather than only 5 values that are assigned in the original
scheme
The pile unit side friction (fp) is obtained from qE and the CPT material index Ic using the following
expression at each elevation along the sides of the pile
10(0732 middot 119868119888 minus 3605) fp middot middot C13 = 119902119864 middot 120579119875119879 middot 120579119879119862 120579119877119860119879119864
C-20
Figure C7 Nine-zone soil behavioral soil type using normalized piezocone parameters
(after Robertson 2009 Mayne 2014)
where θPT = coefficient for pile type (θPT = 084 for bored piles 102 for jacked piles 113 for driven
piles) θTC = coefficient for loading direction (θTC = 111 for compression and 085 for tension) and θRATE
= rate coefficient applied to soils in SBT zones 1 through 7 (θRATE = 109 for constant rate of penetration
tests and 097 for maintained load tests) Since CPT provides data at regular intervals of 2 cm to 5 cm
along the sides of the pile the average fp from z = 0 to z = L can be used directly in Equation 1 to obtain
the shaft capacity Qs
C-21
The pile end bearing resistance is obtained from
middot 10(0325middot 119868119888minus1218) 119902119887 = 119902119864 C14
where qE is averaged in the vicinity of the pile tip Figure C8 shows the Modified Unicone Method in
graphical format For sensitive clays of zone 1 please see additional discussions by Niazi amp Mayne
(2016)
Figure C8 Summary of Modified UniCone Method for direct CPT assessment of axial pile
capacity
C-22
C14 Axial Pile Displacement
The movement of pile foundations can be assessed using elastic continuum theory (Poulos amp
Davis 1980 Randolph 2003 Mayne amp Niazi 2017) These relationships have been developed
using finite element analyses boundary elements and analytical closed-form solutions For the
latter approach the top displacement of a rigid pile subjected to a vertical force is shown in
Figure C9 This gives the movement at the top of a pile subjected to either compression or
tension (uplift) loading Also the percentage of axial load transferred to the pile toe is
determined For piles extending through various soil layers the elastic solution can be
implemented by using a set of stacked pile segments each with its own stiffness as
represented by a soil Youngs modulus
For pile groups the use of computer software is recommended Several available programs can handle
pile groups under axial and lateral moment loading such as DEFPIG (Univ Sydney) and PIGLET (Univ
Western Australia) A full listing of pile foundation software is given at the Geotechnical amp
GeoEnvironmental Service Directory wwwggsdcom
Ps
= Sh
aft
Load
sL
t
tEd
IPw
Load Transfer to Base
)]1)((5ln[
)(
)1(1
1
2 vdL
dLFI
E
ECT
211
IF
P
P CT
t
b
Base Load = Pb
Randolph Elastic Solution
Ground Surface
Top Load Pt = Ps + Pb
Rigid Pile Response
wt = displacementd = diameter
L = length
Es = Elastic Soil Modulus
Es0 (at z = 0) surface
EsM (at z = L2)mid-length
EsL (at z = L)full length
Load Transfer to Shaft
E = EsMEsL = Gibson parameterE = 1 for homogeneous case (constant E)E = 05 for pure Gibson case (Es0 = 0)
where FCT = load direction (= 1 compression 0 tension)
21
IF
P
P CT
t
b
z = depth
= Poissons ratio
Tension
Compression
C-23
Figure C9 Elastic continuum solution for axial pile response under compression and tension
loading
C-24
C15 Acknowledgements
The author appreciates the help of Fernando Illingworth of PonsEcuador Meena Viswanth of
GeoSyntecAtlanta and Dr Marco Uzielli of GeoRiskItaly in helping to collect and process the data
Funding support was provided by David Woeller of ConeTec Richmond BC
C-25
- Structure Bookmarks
-
- CONE PENETRATION TEST DESIGN GUIDE FOR STATE GEOTECHNICAL ENGINEERS
- TABLE OF CONTENTS
- LIST OF TABLES
- CHAPTER 1 INTRODUCTION
- CHAPTER 2 DIRECT CPT METHOD FOR SOIL CHARACTERIZATION
- CHAPTER 3 DIRECT CPT METHOD FOR SHALLOW FOUNDATIONS
- CHAPTER 4 DIRECT CPT METHOD FOR DEEP FOUNDATIONS
- REFERENCES
- APPENDIX A
- APPENDIX B
- APPENDIX C
-
Technical Report Documentation Page
1 Report No
MNRC 2018-32
2 3 Recipients Accession No
4 Title and Subtitle
Cone Penetration Test Design Guide for State Geotechnical
Engineers
5 Report Date
November 2018 6
7 Author(s)
Ryan Dagger David Saftner and Paul Mayne
8 Performing Organization Report No
9 Performing Organization Name and Address
Department of Civil Engineering University of Minnesota Duluth 1405 University Drive Duluth MN 55812
10 ProjectTaskWork Unit No
CTS2017022 11 Contract (C) or Grant (G) No
(C) 99008 (WO) 249
12 Sponsoring Organization Name and Address
Minnesota Department of Transportation Research Services amp Library 395 John Ireland Boulevard MS 330 St Paul Minnesota 55155-1899
13 Type of Report and Period Covered
Final Report 14 Sponsoring Agency Code
15 Supplementary Notes
httpmndotgovresearchreports2018201832pdf 16 Abstract (Limit 250 words)
The objectives of this project are focused on a new cone penetration testing (CPT) geotechnical design manual for highway and transportation applications based on recent research and innovation covering the period from 2000 to 2018 A step-by-step procedure is outlined on how to use CPT data in the analysis and design of common geotechnical tasks Previous manuals are either very outdated with information from 1970-1996 or not appropriately targeted to transportation works This design document introduces modern and recent advancements in CPT research not otherwise captured in legacy manuals from the 1990rsquos and earlier Examples and case studies are provided for each topic interpreted using CPT measures In the manual a step-by-step procedure is outlined on how to use CPT data in analysis and design for typical geotechnical practices These topics which are applicable both to state highways and local roads include bridge foundations (including shallow footings and deep foundations) and soil characterization (including determination of standard soil engineering properties)
17 Document AnalysisDescriptors
Cone penetrometers Geographic information systems Soils by
properties Bridge foundations Soil mechanics
18 Availability Statement
No restrictions Document available from
National Technical Information Services
Alexandria Virginia 22312
19 Security Class (this report)
Unclassified
20 Security Class (this page)
Unclassified
21 No of Pages
225
22 Price
CONE PENETRATION TEST DESIGN GUIDE FOR STATE
GEOTECHNICAL ENGINEERS
Prepared by
Ryan Dagger
David Saftner
Department of Civil Engineering
University of Minnesota Duluth
Paul W Mayne
School of Civil amp Environmental Engineering
Georgia Institute of Technology Atlanta
November 2018
Published by
Minnesota Department of Transportation
Research Services amp Library
395 John Ireland Boulevard MS 330
St Paul Minnesota 55155-1899
This report represents the results of research conducted by the authors and does not necessarily represent the views or policies
of the Minnesota Department of Transportation the University of Minnesota Duluth or the Georgia Institute of Technology
This report does not contain a standard or specified technique
The authors the Minnesota Department of Transportation the University of Minnesota Duluth and the Georgia Institute of
Technology do not endorse products or manufacturers Trade or manufacturersrsquo names appear herein solely because they are
considered essential to this report because they are considered essential to this report
TABLE OF CONTENTS
CHAPTER 1 INTRODUCTION1
CHAPTER 2 DIRECT CPT METHOD FOR SOIL CHARACTERIZATION 4
21 Introduction 4
22 Soil Unit Weight 6
23 CPT Material Index 7
231 Step 1 Normalized Sleeve Friction 7
232 Step 2 Iteration 8
24 Soil Behavior Type (SBT) 8
241 Step 1 8
242 Step 2 9
25 Effective Stress Friction Angle 9
26 Stress History 10
27 Lateral Stress Coefficient 11
28 Undrained Shear Strength 12
29 Ground Stiffness and Soil Moduli 12
210 Coefficient of Consolidation 13
211 Hydraulic Conductivity14
212 Example Problems 15
2121 Example 1 Direct CPT Methods for Geoparameters on Sands 15
2122 Example 2 Direct CPT Methods for Geoparameters on Clay 33
CHAPTER 3 DIRECT CPT METHOD FOR SHALLOW FOUNDATIONS 56
31 Procedure 56
311 Step 1 Estimating Footing Dimensions57
312 Step 2 Soil Characterization 58
313 Step 2a Foundation soil formation parameter59
314 Step 3 Soil elastic modulus and Poissonrsquos ratio 60
315 Step 4 Net cone tip resistance 60
316 Step 5 Bearing capacity of the soil 61
317 Step 6 Settlement61
318 Step 7 Final Check 61
32 Example Problems 62
321 Example 3 Direct CPT Method on Sands 62
CHAPTER 4 DIRECT CPT METHOD FOR DEEP FOUNDATIONS 70
41 Introduction70
42 Modified UniCone Method72
421 Step 172
422 Step 272
423 Step 373
424 Step 473
43 Axial Pile Displacements 73
44 Example Problems 74
441 Example 5 Direct CPT Methods Axial Pile Capacity74
REFERENCES 87
APPENDIX A
APPENDIX B
APPENDIX C
1
2
3
4
5
6
7
8
9
10
11
12
13
14
15
16
17
18
19
20
21
22
23
24
LIST OF FIGURES
Figure Geoparameters determined from CPT 1
Figure Conventional method for shallow foundation design compared to direct CPT method 2
Figure Direct CPT evaluation of axial pile capacity 3
Figure Soil unit weight from CPT sleeve friction 6
Figure Approximate ldquorules of thumbrdquo method using CPT sounding from Wakota Bridge MN 7
Figure SBT zones using CPT Ic 9
Figure Yield stress exponent compared to CPT material index 11
Figure k vs dissipation time for 50 consolidation (Mayne 2017) 15
Figure CPT data from Benton County Minnesota for example problem 1 16
Figure Soil layers using ldquorules of thumbrdquo17
Figure Soil layer 1 using SBT method 23
Figure Soil layer 2 using SBT method 24
Figure Soil layer 3 using SBT method 25
Figure Soil layer 4 using SBT method 26
Figure CPT data from Minnesota for example problem 2 34
Figure Soil layers using ldquorules of thumbrdquo36
Figure Soil layer 1 using SBT method 42
Figure Soil layer 2 using SBT method 43
Figure Soil layer 3 using SBT method 44
Figure Soil layer 4 using SBT method 45
Figure Dissipation t50 data 53
Figure Direct CPT method for shallow foundations56
Figure Direct CPT method introduction 58
Figure Differentiation of porewater pressure measurement locations (Lunne et al 1997) 58
Figure 25 Foundation soil formation parameter hs versus CPT material index Ic (Mayne 2017) 60
Figure 26 Diagram of footing profiles for Example 362
Figure 27 CPT data from Northern Minnesota 63
Figure 28 Schematic of foundation design associated with CPT data 64
Figure 29 Soil type for example problem 367
Figure 30 Direct CPT evaluation of axial pile capacity 70
Figure 31 UniCone Method soil behavior type using CPT (Mayne 2017) 71
Figure 32 Modified UniCone Method soil behavior type using CPT (Mayne 2017) 71
Figure 33 Deep foundation end bearing pile diagram74
Figure 34 CPT data from Minnesota for example problem 5 75
Figure 35 Soil layers using ldquorules of thumbrdquo for pile capacity example using direct CPT method 76
Figure 36 Soil layer 1 using SBT method 81
Figure 37 Soil layer 2 using SBT method 82
Figure 38 Soil layer 3 using SBT method 83
Figure 39 Soil layering compared to pilings 85
LIST OF TABLES
Table 1 Geoparameters calculated directly from CPT 4
Table 2 Minor Geoparameters 5
Table 3 Soil type compared to exponent m 10
Table 4 Geoparameters evaluated for Case Example 117
Table 5 Geoparameters 35
Table 6 Bearing capacity defined by soil type61
Table 7 SCPT Results 62
CHAPTER 1 INTRODUCTION
The purpose of this manual is to provide an analysis of soil characterization shallow footings and deep
foundations using direct cone penetration testing (CPT) methods Geotechnical site characterization is
important for evaluating soil parameters that will be used in the analysis and design of foundations
retaining walls embankments and situations involving slope stability Common practice is to determine
these soil parameters through conventional lab and in-situ testing An alternative method uses CPT
readings of cone tip resistance (qt) sleeve friction ( fs) and pore pressure (u2) directly to determine these
parameters such as unit weight effective friction angle undrained shear strength and many others
(Figure 1) Direct CPT methods are provided for these parameters which will be used in the designs for
shallow and deep foundations
Figure 1 Geoparameters determined from CPT
Designing shallow foundations is typically done in a two-part process determining the bearing capacity
and expected settlement (commonly referred to as displacement) of the soil to approximate the
required size and shape of a foundation The older traditional methods are no longer required with
many approaches existing for using CPT directly in the design of shallow foundations Results from these
methods can provide a direct assessment of bearing capacity andor settlement A specific approach to
the direct method has been recommended and tested using a database of 166 full-scale field load tests
(Figure 2) A one-part process is used to scale the measured cone penetrometer readings (ie
measured cone tip resistance sleeve friction and pore water pressure) to obtain the bearing capacity
and settlement of the soil A step-by-step procedure has been created to transition from the CPT data to
bearing capacity with settlement accounted for The steps consist of estimating a design footing width
1
and length while using the process in the Soil Characterization section to determine soil parameters
directly from the CPT data
Figure 2 Conventional method for shallow foundation design compared to direct CPT method
There are upwards of 40 different direct CPT methods that have been developed over the past five
decades to determine the axial compression capacity of a piling foundation Earlier direct CPT methods
relied on hand-recorded information where mechanical-type CPT cone tip resistance data would be
collected at 20 cm intervals
Herein the method recommended for deep foundation design is the Modified UniCone method which
uses all three readings of the modern electronic piezocone penetrometer (CPTu) while addressing a
variety of pile foundation types (Figure 3) The modified UniCone Method is based on a total of 330 pile
load tests (three times the original UniCone database) that were associated with SCPTu data Two
computer software programs have been recommended for analyzing pile movements andor
settlements
2
Figure 3 Direct CPT evaluation of axial pile capacity
3
Symbol Parameter
γt Soil total unit weight
Ic CPT material index
SBT Soil behavior type (SBT)
ʹ σp Preconsolidation stress
YSR Yield stress ratio
ϕʹ Effective friction angle
Ko Lateral stress coefficient
su Undrained shear strength
Dʹ Constrained modulus
Eʹ Drained Youngrsquos modulus
CHAPTER 2 DIRECT CPT METHOD FOR SOIL
CHARACTERIZATION
21 INTRODUCTION
A direct CPT method for determining the value of each of various geoparameters shown in Table 1 is
provided The parameter Ic is used in the derivation of several sequential parameters This section of
the guide may be referred to as these parameters are used in calculations throughout the shallow
foundations and deep foundations sections
Table 1 Geoparameters calculated directly from CPT
4
Kʹ Bulk modulus
ks Subgrade reaction modulus
MR Resilient modulus
Gmax Small-strain shear modulus
k Coefficient of permeability
cv Coefficient of consolidation
Symbol Parameter Equation
ρt Mass Density ρt = γtga
where ga = 98 ms2
σvo Total Stress σvo asymp Σ (γti middot Δzi)
c Effective cohesion ʹ Empirical c asymp 003σp
In clays c asymp 01cu
Other minor parameters can be used in calculations of the geoparameters in Table 1 These minor parameters are provided in Table 2
Table 2 Minor Geoparameters
5
22 SOIL UNIT WEIGHT
The total soil unit weight can be estimated from CPT sleeve friction resistance as shown in Figure 4
(Mayne 2014) This method is not applicable to organic clays diatomaceous soils peats or sensitive
soils
Figure 4 Soil unit weight from CPT sleeve friction
Use Equation 1 to calculate the soils total unit weight
1 120574119905 = 120574119908 ∙ [122 + 015 ∙ ln (100 ∙ 119891119904 + 001)]
120590119886119905119898
Since soil unit weight is required for determining most geoparameters (including Soil Behavior Type)
estimating the soil type of the layers using the ldquorules of thumbrdquo method is a good first step before
determining more precise layering (Figure 5) Once a unit weight is determined for each noticeable layer
change these results can be used in later calculations such as ldquoCPT Material Indexrdquo To use the ldquorules
of thumbrdquo method some helpful guidelines are to assume sands are identified when qt gt 725 psi and u2
asymp uo while the presence of intact clays are prevalent when qt lt 725 psi and u2 gt uo The magnitude of
porewater pressures help to indicate intact clays such as soft (u2 asymp 2middotuo) firm (u2 asymp 4middotuo) stiff (u2 asymp 8middotuo)
and hard (u2 asymp 20middotuo) Fissured overconsolidated clays tend to have negative u2 values such that u2 lt 0
6
Figure 5 Approximate ldquorules of thumbrdquo method using CPT sounding from Wakota Bridge MN
23 CPT MATERIAL INDEX
The development of the CPT material index Ic has improved the initial classification of soil types and
calculation of soil parameters as shown in Table 1 To calculate Ic follow steps 1 and 2
231 Step 1 Normalized Sleeve Friction
Calculate Fr using Equation 2 if not provided with the CPT data gathered
2
7
119891119904 119865119903() = 100 ∙
(119902119905minus120590119907119900)
232 Step 2 Iteration
Iterate using Equations 3-5 to determine Ic by initially using n = 1 to calculate a starting value of Ic The
exponent n is soil-type dependent n = 1 (clays) n asymp 075 (silts) and n asymp 05 (sands) Iteration converges
quickly which is generally after the 3rd cycle
3 (119954119957minus120648119959119952)120648119938119957119950 = 119951 119928119957119951 prime (120648119959119952120648119938119957119950)
4
prime 120590119907119900 119899 = 0381 ∙ 119868119888 + 005 ( ) minus 015 120590119886119905119898
119899 le 10
5 119920119940 = radic[120785 120786120789 minus 119949119952119944119928119957119951]120784 + [120783 120784120784 + 119949119952119944119917119955]120784
24 SOIL BEHAVIOR TYPE (SBT)
Typically soil samples are not taken when using CPT The soil types are then inferred from the qt fs and
u2 readings To determine the types of soil from CPT data follow steps 1 and 2
241 Step 1
To determine the soil layers from CPT results calculate Ic by following the steps under section ldquoCPT
Material Indexrdquo After Ic has been determined through all specified depths use Figure 6 to classify the
type of soil by comparing each Ic value to normalized CPT readings (Fr and Qtn) from Equations 2 and 3
8
Figure 6 SBT zones using CPT Ic
242 Step 2
To determine if any of the soil layers contain ldquosensitive clays and siltsrdquo from zone 1 or ldquovery stiff
overconsolidated (OC) soilrdquo from zones 8 and 9 use Equations 6 and 7 If any soil layers are found within
zone 1 by Equation 6 then caution should be taken as these clays are prone to instability collapses and
difficulties in construction performance Very stiff OC sands to clayey sands of zone 8 (15 lt Fr lt 45)
and very stiff OC clays to silts of zone 9 (Fr gt 45) can be identified by Equation 7
6 Equation 6 errata This is an exponential expression (see Figure 5)
119876119905119899 lt 12119890119909119901(minus14 ∙ 119865119903)
7 119876119905119899 gt 0005(119865119903minus1)minus00003(119865119903minus1)2minus0002
1
Errata See terms and coefficients in Figure 5 above (numbers cannot be rounded off)
After each CPT reading has been assigned a zone from Figure 14 a visual representation can be made to
show the predominant layers by soil types (Figure A5)
25 EFFECTIVE STRESS FRICTION ANGLE
The effective friction angle (ϕ) is used to govern the strength for sands and clays where Equation 8 is
used for sands and Equation 9 is used for clays The value of ϕ for sands is derived from Ic so before ϕ
can be calculated refer to the iteration of Qtn under the ldquoCPT Material Indexrdquo section Once the values
9
Soil Type m
Fissured clays 11
Intact clays 10
Sensitive clays 09
Silt mixtures 085
Silty sands 080
Clean sands 072
Note m may be higher than 11 in fissured clays
of Ic and Qtn are calculated the type of soil can be determined from the section ldquoSoil Behavior Type
SBTrdquo The type of soil will dictate which equation to use for ϕ
Sands
8 120601prime(deg) = 176deg + 110deg log(119876119905119899)
Clays
9 119902119905minus120590119907119900 120601prime(deg) = 295deg ∙ 119861119902
0121 ∙ [0256 + 0336 ∙ 119861119902 + log ( )] prime 120590119907119900
Where = 119861119902 (1199062 minus 119906119900)(119902119905 minus 120590119907119900)
26 STRESS HISTORY
Determining the stress history can be characterized by an apparent yield stress ratio of the form
10 prime 120590119901
119884119878119877 = prime 120590119907119900
Where σp is defined as the preconsolidation stress or effective yield stress (Equation 10) The YSR is the
same equation as the more common overconsolidation ratio (OCR) but is now generalized to
accommodate mechanisms of preconsolidation such as ageing desiccation repeated cycles of wetting-
drying repeated freeze thaw cycles and other factors
11 prime prime 120590119901 = 120590119910 = 033(119902119905 minus 120590119907119900)119898prime
Where σp σy σvo and qt have units of kPa and the value of m depends on soil type with typical values shown in Table 3
Table 3 Soil type compared to exponent m
10
The value of m for non-fissured soils and inorganic clays and silts is derived from Ic (Figure 7) so before
m can be calculated (Equation 12) refer to the iteration of Ic under the ldquoCPT Material Indexrdquo section
Determine Ic for all soil layers before calculating m
12 028
119898prime = 1 minus 25 119868119888 1+( frasl ) 265
Once m is known for all soil layers the YSR can be determined
Figure 7 Yield stress exponent compared to CPT material index
A limiting value of YSR can be reached for clays and sands It can be calculated using Equation 13
13 (1 sin(emptyprime))
(1+sin(emptyprime)) 119884119878119877119897119894119898119894119905 = [ ]
(1minussin(emptyprime))2
27 LATERAL STRESS COEFFICIENT
The lateral stress coefficient Ko = σhoσvo commonly referred to as the at-rest condition is used to
represent the horizontal geostatic state of soil stress Ko can be calculated using Equation 14 but
11
14
119870119900 = (1 minus sin(emptyprime)) ∙ 119884119878119877sin( emptyprime)
sections such as ldquoStress Historyrdquo and ldquoEffective Stress Friction Anglerdquo will need to be referred to for
determining parameters ϕ and YSR
A maximum value for Ko can be determined by Equation 15
15 (1+sin(emptyprime))
119870119900119898119886119909 = = 1199051198861198992(45deg + emptyprime2) (1minussin(emptyprime))
28 UNDRAINED SHEAR STRENGTH
Loading on soils can result in fully drained partially drained or fully undrained conditions Sands
typically produce drained cases due to their high permeability but exceptions may occur in loose sands
during fast loading where the water does not have sufficient time to dissipate Clays exhibit low
permeability and thus often result in undrained loading cases when a load is applied quickly For soft-
firm clays the undrained shear strength (su) can be determined from CPT via Equation 16 where the
value of the bearing factor Nkt can be taken as 12
16 119902119905minus120590119907119900 119904119906 =
119873119896119905
In the case of remolded undrained shear strength from CPT su asymp fs
29 GROUND STIFFNESS AND SOIL MODULI
Determining the grounds stiffness can be measured from geoparameters such as the constrained
modulus (D) drained Youngrsquos modulus (E) bulk modulus (K) subgrade reaction modulus (ks) resilient
modulus (MR) and small-strain shear modulus (Gmax)
17 119863 prime asymp 5 ∙ (119902119905 minus 120590119907119900)
18
12
119864 prime = 119863 prime
11
19 119864prime
119870prime = [3∙(1minus2119907prime)]
The subgrade modulus (ks) is a combination of soil-structural properties which creates a parameter that
depends on the ground stiffness and the size of the loaded element
20 119864prime
119896119904 = [119889∙(1minus1199072)]
The resilient modulus MR applies to pavement analysis and design and can be calculated using Equation
21 where qt and fs are in MPa
21 053 119872119877 = (146119902119905 + 135511989111990414 + 236)244
The small strain shear modulus Gmax is a representation of the initial stiffness of all soils and rocks
Graphically it is the beginning portion of all stress-strain-strength curves for geomaterials Use Equation
22 to determine Gmax
22 119866119898119886119909 = 120588119905 ∙ 1198811199042
Where Vs = shear wave velocity (Equation 23) as measured by seismic cone penetration tests (SCPT) If
only cone penetration tests (CPT) or piezocone (CPTu) data are available the shear wave velocity may
be estimated from
23 03 119891119904 119881119904 (119898frasl119904) = [101 ∙ log(119902119905) minus 114]167 ∙ (100 ∙ )
119902119905
Where qt and fs have units of (kPa)
210 COEFFICIENT OF CONSOLIDATION
The coefficient of consolidation (cv) controls the rate that foundation and embankment settlements
occur By using results of CPT dissipation tests that measure the change in u2 readings over time the
value of cv can be determined Using Equation 24 cv can be determined based on piezocone dissipation
13
curves The equation below requires an estimate of the in-situ rigidity index (IR) of the soil If results of
SCPTU are available then IR may be determined from Gmax and qt per equation A38
24 0030∙(119886119888)2∙(119868119877)075
119888119907 = 11990550
Where ac = penetrometer radius (178 cm for 10-cm2 cone 220 cm for 15-cm2 cone) t50 = time to reach 50 dissipation IR = Gsu = undrained rigidity index G can be determined from the ldquoGround Stiffness and Soil Modulirdquo section
211 HYDRAULIC CONDUCTIVITY
The hydraulic conductivity also known as the coefficient of permeability (k) expresses the flow
characteristics of soils and has units of cms or feetday One method of calculation would be to use
Equation 25 where cv would need to be determined from ldquoCoefficient of Consolidationrdquo and D would
need to be determined from ldquoGround Stiffness and Soil Modulirdquo
25 119888119907∙120574119908 119896 =
119863prime
An alternative approach developed for soft normally-consolidated soils (Figure 8) is shown in Equation
26 where t50 (sec) values are used directly in assessing k in (cms)
26
14
125
119896 asymp ( 1
) 251∙11990550
Figure 8 k vs dissipation time for 50 consolidation (Mayne 2017)
212 EXAMPLE PROBLEMS
2121 Example 1 Direct CPT Methods for Geoparameters on Sands
Several geoparameters need to be determined based on the given CPT data collected for the South
Abutment of a bridge in Benton County MN (Figure 9) The groundwater table (GWT) was measured at
17 feet Determine all the geoparameters found in Table 4 at depths of 0 feet to 30 feet All sand layers
can be assumed ldquodrainedrdquo with ν = 02
15
Figure 9 CPT data from Benton County Minnesota for example problem 1
16
Table 4 Geoparameters evaluated for Case Example 1
Symbol Parameter Symbol Parameter
γt Soil total unit weight Ko Lateral stress coefficient
Ic CPT material index Dʹ Constrained modulus
SBT Soil behavior type (SBT) Eʹ Drained Youngrsquos modulus ʹ σp Preconsolidation stress Kʹ Bulk modulus
YSR Yield stress ratio MR Resilient modulus
ϕʹ Effective friction angle Gmax Small-strain shear modulus
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 10 Soil layers using ldquorules of thumbrdquo
γw = 6224 pcf
17
Layer 1
From Figure 10 fs = 17 psi taken as a representative value of the layer
17 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1204 pcf
145 psi
Layer 2
From Figure 10 fs = 17 psi taken as a representative value of the layer
17psi
γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1204 pcf 145 psi
Layer 3
From Figure 10 fs = 12 psi taken as a representative value of the layer
12 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 4
From Figure 10 fs = 7 psi taken as a representative value of the layer
CPT Material Index
7 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1121 pcf
145 psi
Layer 1
From Figure 10 fs = 17 psi and qc = 3500 psi
18
qt = qc + u2 ∙ (1 minus a) = 3500 psi + 3 psi ∙ (1 minus 08) = 3501 psi
= γt ∙ 6 feet = 1204 pcf ∙ 6 feet = 7225 psf = 50 psi σvo
fs 17 psi Fr() = 100 ∙ = 100 ∙ = 049
(qt minus σvo) (3501 psi minus 50 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 5 psi minus 0 psi = 50 psi
(qt minus σvo)σatm (3501 psi minus 50 psi )145 psi = = Qtn prime )n (σvoσatm (50 psi145 psi)n
prime σvo 50 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(049)]2
Qtn = 35329 n = 036 lt 10 Ic = 13
Layer 2
From Figure 10 fs = 17 psi and qc = 3500 psi
19
qt = qc + u2 ∙ (1 minus a) = 3500 psi + 0 psi ∙ (1 minus 08) = 3500 psi
= 7225 psf + γt ∙ 6 feet = 7225 psf + 1204 pcf ∙ 6 feet = 1445 psf = 100 psi σvo
fs 17 psi Fr() = 100 ∙ = 100 ∙ = 049
(qt minus σvo) (3500 psi minus 100 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 100 psi minus 0 psi = 100 psi
(qt minus σvo)σatm (3500 psi minus 100 psi )145 psi = = Qtn prime )n (σvoσatm (100 psi145 psi)n
prime σvo 100 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
47 minus log(Qtn)]2 + [122 + log(049)]2
Ic = radic[3
Qtn = 27947 n = 041 lt 10 Ic = 14
Layer 3
From Figure 10 fs = 12 psi and qc = 1500 psi
20
qt = qc + u2 ∙ (1 minus a) = 1500 psi + 0 psi ∙ (1 minus 08) = 1500 psi
= 1445 psf + γt ∙ 11 feet = 1445 psf + 1172 pcf ∙ 11 feet = 2734 psf = 190 psi σvo
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 081
(qt minus σvo) (1500 psi minus 190 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = γwater ∙ (z minus 119911119908) = 6224 pcf ∙ (23 feet minus 17 feet) = 3734 psf = 99 psi
prime σvo = σvo minus uo = 190 psi minus 99 psi = 164 psi
(qt minus σvo)σatm (1500 psi minus 190 psi )145 psi = = Qtn prime )n (σvoσatm (164 psi145 psi)n
prime σvo 164 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(081)]2
Qtn = 947 n = 06 lt 10 Ic = 19
21
Layer 4
From Figure 10 fs = 7 psi and qc = 1200 psi
qt = qc + u2 ∙ (1 minus a) = 1200 psi + 0 psi ∙ (1 minus 08) = 1200 psi
= 2734 psf + γt ∙ 11 feet = 2734 psf + 1121 pcf ∙ 7 feet = 3519 psf = 244 psi σvo
fs 7 psi Fr() = 100 ∙ = 100 ∙ = 060
(qt minus σvo) (1200 psi minus 244 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = γwater ∙ (z minus 119911119908) = 6224 pcf ∙ (30 feet minus 17 feet) = 8091 psf = 130 psi
prime σvo = σvo minus uo = 244 psi minus 130 psi = 188 psi
(qt minus σvo)σatm (1200 psi minus 244 psi )145 psi = = Qtn prime )n (σvoσatm (188 psi145 psi)n
prime σvo 188 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(060)]2
22
Qtn = 686 n = 064 lt 10 Ic = 19
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic Qtn and Fr the first layer is defined as a ldquoDrained Gravelly Sandrdquo from
Figure 11
Figure 11 Soil layer 1 using SBT method
23
Layer 2
Based on values of Ic Qtn and Fr the second layer is defined as a ldquoDrained Sandrdquo from Figure
12
Figure 12 Soil layer 2 using SBT method
24
Layer 3
Based on values of Ic Qtn and Fr the third layer is defined as a ldquoDrained Sandrdquo from Figure 13
Figure 13 Soil layer 3 using SBT method
25
Layer 4
Based on values of Ic Qtn and Fr the fourth layer is defined as a ldquoDrained Sandrdquo from Figure 14
Figure 14 Soil layer 4 using SBT method
Effective Stress Friction Angle
Layer 1
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(3533) = 456deg
Layer 2
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(2795) = 445deg
26
Layer 3
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(947) = 393deg
Layer 4
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(686) = 378deg
Stress History
Layer 1
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (13frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(24139 kPa minus 345 kPa)072 = 4717 kPa
= 684 psi
prime σp 684 psi YSR = = = 136
primeσvo 50 psi
Layer 2
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + (14 1 + ( frasl ) frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(34132 kPa minus 689 kPa)072 = 605 kPa
= 877 psi
prime σp 877 psi YSR = = = 87
primeσvo 100 psi
27
Layer 3
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + (19 1 + ( frasl ) frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(10342 kPa minus 131 kPa)072 = 254 kPa
= 368 psi
prime σp 368 psi YSR = = = 22
primeσvo 164 psi
Layer 4
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (19frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(8274 kPa minus 168 kPa)072 = 215 kPa = 312 psi
prime σp 312 psi YSR = = = 17
primeσvo 188 psi
Lateral Stress Coefficient
Layer 1
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(456deg)) ∙ 136sin( 456deg) = 18
Layer 2
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(445deg)) ∙ 87sin( 445deg) = 14
28
Dprime 17480 psi
Eprime = = = 15890 psi 11 11
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(393deg)) ∙ 22sin( 393deg) = 06
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(378deg)) ∙ 17sin( 378deg) = 05
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3501 psi minus 50 psi) = 17480 psi
Eprime 15890 psi
Kprime = = = 8828 psi [3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Layer 3
Layer 4
Ground Stiffness and Soil Moduli
Layer 1
Values of qt and fs need to be in MPa
MR = (146(241 MPa)053 + 1355(012 MPa)14 + 236)244 = 3416 MPa = 49575 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1172 kPa m Vs (mfrasls) = [101 ∙ log(24139 kPa) minus 114]167 ∙ (100 ∙ ) = 2745
24139 kPa s
Vs = 9012 fts
29
γt 1204 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 37 ∙ (9012 ) = 304 ∙ 106psf = 21000 psi s
Layer 2
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3500 psi minus 100 psi) = 17480 psi
Dprime 17480 psi Eprime = = = 15890 psi
11 11
Eprime 15890 psi Kprime = = = 8828 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(241 MPa)053 + 1355(012 MPa)14 + 236)244 = 3415 MPa = 49562 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1172 kPa m Vs (mfrasls) = [101 ∙ log(24133 kPa) minus 114]167 ∙ (100 ∙ ) = 2745
24133 kPa s
Vs = 9012 fts
γt 1204 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
30
2 ft Gmax = ρt ∙ Vs
2 = 37 ∙ (9012 ) = 304 ∙ 106psf = 21000psi s
Layer 3
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (1500 psi minus 190 psi) = 7405 psi
Dprime 7405 psi Eprime = = = 6732 psi
11 11
Eprime 6732 psi Kprime = = = 3740 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(103 MPa)053 + 1355(008 MPa)14 + 236)244 = 1506 MPa = 21854 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 827 kPa m Vs (mfrasls) = [101 ∙ log(10343 kPa) minus 114]167 ∙ (100 ∙ ) = 2611
10343 kPa s
Vs = 8566 fts
γt 1172 pcf slug ρt = = = 36
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 36 ∙ (8566 ) = 27 ∙ 106psf = 19000 psi s
Layer 4
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (1200 psi minus 244 psi) = 5878 psi
31
Dprime 5878 psi Eprime = = = 5343 psi
11 11
Eprime 5343 psi Kprime = = = 2969 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(83 MPa)053 + 1355(005 MPa)14 + 236)244 = 165 MPa = 16901 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 483 kPa m Vs (mfrasls) = [101 ∙ log(8274 kPa) minus 114]167 ∙ (100 ∙ ) = 2243
8274 kPa s
Vs = 7360 fts
γt 1121 pcf slug ρt = = = 35
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 35 ∙ (7360 ) = 389 ∙ 106psf = 13000 psi s
32
2122 Example 2 Direct CPT Methods for Geoparameters on Clay
Several geoparameters need to be determined based on the given CPT data collected for the South
Abutment (Figure 15) The groundwater table (GWT) was measured at 60 feet Determine all the
geoparameters found in Table 5 at depths of 0 feet to 42 feet All sand layers can be assumed ldquodrainedrdquo
ν = 02 All clay layers can assume ν = 049
33
Figure 15 CPT data from Minnesota for example problem 2
34
Table 5 Geoparameters
Symbol Parameter Symbol Parameter
γt Soil total unit weight Eʹ Drained Youngrsquos modulus
Ic CPT material index Kʹ Bulk modulus
SBT Soil behavior type (SBT) MR Resilient modulus
ʹ σp Preconsolidation stress Gmax Small-strain shear modulus
YSR Yield stress ratio su Undrained shear strength
ϕʹ Effective friction angle cv Coefficient of consolidation
Ko Lateral stress coefficient k Hydraulic conductivity
Dʹ Constrained modulus
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
35
Figure 16 Soil layers using ldquorules of thumbrdquo
Unit weight of water γw = 6224 pcf
Layer 1
From Figure 16 fs = 13 psi taken as a representative value of the layer
13 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1179 pcf
145 psi
Layer 2
From Figure 16 fs = 12 psi taken as a representative value of the layer
12 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 3
From Figure 16 fs = 20 psi taken as a representative value of the layer
36
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
Layer 4
From Figure 16 fs = 2 psi taken as a representative value of the layer
2 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1004 pcf
145 psi
CPT Material Index
Layer 1
From Figure 10 fs = 13 psi and qc = 3000 psi
qt = qc + u2 ∙ (1 minus a) = 3000 psi + 0 psi ∙ (1 minus 08) = 3000 psi
σvo = γt ∙ 2 feet = 1179 pcf ∙ 2 feet = 235 psf = 16 psi
fs 13 psi Fr() = 100 ∙ = 100 ∙ = 043
(qt minus σvo) (3000 psi minus 16 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 16 psi minus 0 psi = 16 psi
(qt minus σvo)σatm (3000 psi minus 16 psi )145 psi = = Qtn prime )n (16 psi145 psi)n (σvoσatm
prime σvo 16 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(043)]2
37
Qtn = 4126 n = 032 lt 10 Ic = 121 therefore sand
Layer 2
From Figure 10 fs = 12 psi and qc = 250 psi
qt = qc + u2 ∙ (1 minus a) = 250 psi + 10 psi ∙ (1 minus 08) = 252 psi
σvo = 235 psf + γt ∙ 30 feet = 235 psf + 1172 pcf ∙ 30 feet = 3751 psf = 260 psi
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 53
(q minus σ ) (252 psi minus 260 psi) t vo
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 260 psi minus 0 psi = 260 psi
(qt minus σvo)σatm (250 psi minus 260 psi )145 psi = = Qtn prime (σvoσatm)n (260 psi145 psi)n
38
prime σvo 260 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(53)]2
Qtn = 87 n = 10 le 10 Ic = 32 therefore clay
Layer 3
From Figure 10 fs = 20 psi and qc = 4000 psi
qt = qc + u2 ∙ (1 minus a) = 4000 psi + 8 psi ∙ (1 minus 08) = 40016 psi
σvo = 3751 psf + γt ∙ 2 feet = 3751 psf + 1219 pcf ∙ 2 feet = 3995 psf = 277 psi
fs 16 psi Fr() = 100 ∙ = 100 ∙ = 054
(qt minus σvo) (30016 psi minus 277 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 277 psi minus 0 psi = 277 psi
(qt minus σvo)σatm (30016 minus 277)145 psi = = Qtn prime (σvoσatm)n (277 psi145 psi)n
39
primeσvo 277 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(054)]2
Qtn = 1962 n = 052 le 10 Ic = 15 therefore sand
Layer 4
From Figure 10 fs = 2 psi and qc = 250 psi
qt = qc + u2 ∙ (1 minus a) = 250 psi + 0 psi ∙ (1 minus 08) = 250 psi
= 3994 psf + γt ∙ 8 feet = 3994 psf + 1004 pcf ∙ 8 feet = 4798 psf = 333 psi σvo
fs 2 psi Fr() = 100 ∙ = 100 ∙ = 092
(qt minus σvo) (250 psi minus 333 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 333 psi minus 0 psi = 333 psi
(qt minus σvo)σatm (250 psi minus 333 psi )145 psi = = Qtn prime (σvoσatm)n (333 psi145 psi)n
40
prime σvo 333 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(092)]2
Qtn = 65 n = 10 lt 10 Ic = 29 therefore clayey silt
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic Qtn and Fr the first layer is defined as a ldquoDrained Sandrdquo from Figure 17
41
Figure 17 Soil layer 1 using SBT method
Layer 2
Based on values of Ic Qtn and Fr the second layer is defined as a ldquoUndrained Clayrdquo from Figure
18
42
Figure 18 Soil layer 2 using SBT method
43
Layer 3
Based on values of Ic Qtn and Fr the third layer is defined as a ldquoDrained Sandrdquo from Figure 19
Figure 19 Soil layer 3 using SBT method
44
Layer 4
Based on values of Ic Qtn and Fr the fourth layer is defined as a ldquoUndrained Silty Mixrdquo from
Figure 20
Figure 20 Soil layer 4 using SBT method
Effective Stress Friction Angle
Layer 1 (sand)
45
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(4126) = 464deg
Layer 2 (clay)
qt minus σvo
ϕprime(deg) = 295deg ∙ Bq0121 ∙ [0256 + 0336 ∙ Bq + log ( )]
primeσvo
Where
(u2 minus uo) (10 psi minus 0 psi) Bq = = = 004
(qt minus σvo) (252 psi minus 260 psi)
252 psi minus 260 psi ϕprime(deg) = 295deg ∙ 00440121 ∙ [0256 + 0336 ∙ 0044 + log ( )]
260 psi
ϕprime(deg) = 245deg
Layer 3 (sand)
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(1962) = 428deg
Layer 4 (clay)
qt minus σvo
ϕprime(deg) = 295deg ∙ Bq0121 ∙ [0256 + 0336 ∙ Bq + log ( )]
primeσvo
Where
(u2 minus uo) (5 psi minus 0 psi) Bq = = = 002
(qt minus σvo) (251 psi minus 333 psi)
252 psi minus 333 psi ϕprime(deg) = 295deg ∙ 0020121 ∙ [0256 + 0336 ∙ 002 + log ( )]
333 psi
ϕprime(deg) = 265deg
Stress History
46
Layer 1
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (12frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(3000 psi minus 33)072 = 1051 psi
prime σp 1051 psi YSR = = = 642
primeσvo 16 psi
Layer 2
028 028
prime m = 1 minus = 1 minus = 099 25 25 1 + (
Icfrasl ) 1 + (32frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(252 psi minus 260)099 = 734 psi
prime σp 734 psi YSR = = = 28
primeσvo 260 psi
Layer 3
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + ( frasl ) 1 + (15frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(4002 psi minus 277)072 = 1288 psi
47
prime σp 1288 psi YSR = = = 46
primeσvo 277 psi
Layer 4
028 028
prime m = 1 minus = 1 minus = 097 25 25 Ic 1 + ( frasl ) 1 + (29frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(250 psi minus 333)097 = 626 psi
prime σp 626 psi YSR = = = 19
primeσvo 333 psi
Lateral Stress Coefficient
Layer 1
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(464deg)) ∙ 642sin( 464deg) = 56
Layer 2
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(245deg)) ∙ 28sin( 245deg) = 090
Layer 3
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(428deg)) ∙ 46sin( 428deg) = 091
Layer 4
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(266deg)) ∙ 19sin( 266deg) = 073
48
Ground Stiffness and Soil Moduli
Layer 1
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3000 psi minus 16 psi) = 14991 psi
Dprime 14991 psi Eprime = = = 13629 psi
11 11
Eprime 13629 psi Kprime = = = 7571 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(207 MPa)053 + 1355(009 MPa)14 + 236)244 = 2816 MPa = 40873 psi
fs
03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 896 kPa m Vs (mfrasls) = [101 ∙ log(20685 kPa) minus 114]167 ∙ (100 ∙ ) = 2564
20685 kPa s
Vs = 8412 fts
γt 1179 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
2 ft 2 Gmax = ρt ∙ Vs = 37 ∙ (8412 ) = 260 ∙ 106psf = 17992 psi
s
Layer 2
49
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (252 psi minus 260 psi) = 11298 psi
Dprime 11298 psi Eprime = = = 10271 psi
11 11
Eprime 10271 psi Kprime = = = 17118 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(049))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(17 MPa)053 + 1355(008 MPa)14 + 236)244 = 443 MPa = 64309 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 827 kPa m Vs (mfrasls) = [101 ∙ log(17375 kPa) minus 114]167 ∙ (100 ∙ ) = 2646
17375 kPa s
Vs = 8680 fts
γt 1172 pcf slug ρt = = = 36
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 36 ∙ (8680 ) = 274 ∙ 106psf = 19036 psi s
Layer 3
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (4002 psi minus 277 psi) = 19869 psi
Dprime 19869 psi Eprime = = = 18063 psi
11 11
50
Eprime 18063 psi Kprime = = = 10035 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(276 MPa)053 + 1355(014 MPa)14 + 236)244 = 4017 MPa = 58309 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1379 kPa m Vs (mfrasls) = [101 ∙ log(27591 kPa) minus 1379]167 ∙ (100 ∙ ) = 2854
27591 kPa s
Vs = 9363 fts
γt 121 pcf slug ρt = = = 38
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 38 ∙ (9363 ) = 33 ∙ 106psf = 23050 psi s
Layer 4
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (250 psi minus 333 psi) = 1088 psi
Dprime 1088 psi Eprime = = = 989 psi
11 11
Eprime 989 psi Kprime = = = 16490 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(049))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
51
MR = (146(17 MPa)053 + 1355(001 MPa)14 + 236)244 = 360 MPa = 52189 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 138 kPa m Vs (mfrasls) = [101 ∙ log(17238 kPa) minus 114]167 ∙ (100 ∙ ) = 1545
17238 kPa s
Vs = 5069 fts
γt 1004 pcf slug ρt = = = 31
ga 322 ftfrasls2 ft3
2 ft 2 Gmax = ρt ∙ Vs = 31 ∙ (5069 ) = 80 ∙ 105psf = 5565 psi
s
Undrained Shear Strength
Layer 2
qt minus σv0 250 psi minus 26 psi su = = = 187 psi
12 12
Layer 4
qt minus σv0 250 psi minus 333 psi su = = = 181 psi
12 12
Coefficient of Consolidation
52
Example dissipation data are shown in Figure 21 Example calculations are provided
Figure 21 Dissipation t50 data
Layer 1
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (7197)075
cv = = = 359 cmsec 1 sec t50
Layer 2
53
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (94392)075
cv = = = 0008 cmsec 3000 sec t50
Layer 3
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (8303)075
cv = = = 665 cmsec 06 sec t50
Layer 4
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (26714)075
cv = = = 087 cmsec 11 sec t50
Hydraulic Conductivity
Layer 1
cm 1 psi 359 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 10 ∙ 10minus4 cmsec Dprime 14984 psi
Layer 2
cm 1 psi 0008 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 73 ∙ 10minus7 cmsec Dprime 11379 psi
Layer 3
cm 1 psi 665 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 20 ∙ 10minus5 cmsec Dprime 148650 psi
54
Layer 4
cm 1 psi 087 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 18 ∙ 10minus4 cmsec Dprime 10861 psi
55
CHAPTER 3 DIRECT CPT METHOD FOR SHALLOW
FOUNDATIONS
31 PROCEDURE
Shallow foundation analysis is typically done in a two-part traditional procedure The traditional
techniques are no longer required as a direct CPT method for square rectangular and circular shallow
footings is available (Figure 22) This process has the soil types grouped into four main categories sands
silts fissured clays and intact clays When determining soil types for each design it is believed footings
on sands and silts act in a fully drained manner while intact clays act in an undrained manner under
conditions of constant volume In order to determine the vertical stress-displacement-capacity of
square rectangular and circular shallow footings follow the steps provided towards the solution given
by Equation 27
Figure 22 Direct CPT method for shallow foundations
Equation 27 may be used to calculate all footing stresses from zero to the bearing capacity (qmax) To
calculate qmax for a sized footing of width (B) length (L) and thickness (t) follow steps 1 through 7
provided The settlement (s) can be determined after the calculation of qmax by simply rearranging
56
Equation 27 where the allowable stress (qallow) is defined as qmax divided by the factor of safety (FS) For
shallow footings a FS value of 3 is commonly used in geotechnical engineering
120782120787 minus120782120785120786120787 119956 119923 119954119950119938119961 = 119945119956 ∙ 119954119957119951119942119957 ∙ (119913
) ∙ (119913
) 27 119950119938119961
2 119902119898119886119909 0345 1 frasl119865119878 119871 119904 = 119861 ∙ [ ∙ ∙ ( ) ] 28
119861 ℎ119904 119902119905119899119890119905
Where hs = the foundation soil formation parameter qtnet = the net corrected cone tip resistance
311 Step 1 Estimating Footing Dimensions
In Minnesota frost heave can have devastating effect on a shallow foundations It is common practice to
place a foundation bearing elevation below the expected maximum frost depth (roughly 45 to 6 feet
below ground level) With this assumption estimate a footing size (B x L) for design to obtain
representative data from CPT roughly 15middotB below the foundation depth (Df) as shown in Figure 23 The
CPT data collected will consist of
cone tip resistance (qc)
measured porewater pressure acting behind the cone tip (u2) as shown in Figure 24
sleeve friction (fs)
57
Figure 23 Direct CPT method introduction
Figure 24 Differentiation of porewater pressure measurement locations (Lunne et al 1997)
312 Step 2 Soil Characterization
The soil behavior type (SBT) and CPT material index (Ic) govern the value of the formation factor hs The
first step is following the steps in Soil Unit Weight to determine soil layering and γt values for each
layer To determine a representative unit weight (γsoil) for all the layers in the range of Df to 15∙B below
58
Df the user will need to use their engineering judgement on the definition of ldquorepresentative unit
weightrdquo This single value of unit weight is used in further calculations such as the total vertical soil stress (σvo) from Equation 29 and effective vertical stress (σvo) from Equation 30 Values of σvo and σvo
will need to be calculated at 15middotB below Df
120648119959119952 = sum(120632119956119952119946119949 ∙ 119963) 29
prime 120648119959119952 = 120648119959119952 minus 119958119952 30
Where z = Df +15middotB uo=γwatermiddot(z-zw)
The qc will need to be corrected using Equation 31 in the case of fine-grained soils that develop excess porewater pressure during cone penetration These values will be used in the following calculations
119954119957 = 119954119940 + 119958120784 ∙ (120783 minus 119938) 31
Where 119860119899 a = cone area ratio = eg MnDOT commonly uses a = 08 119860119888
119860119899 = cross-sectional area of load cell or shaft 119860119888 = projected area of the cone
The cone area ratio is determined based on the type of piezocone tip used during in-situ field testing
Manufacturer specifications should provide the measured net area ratio (a) for the particular cone
penetrometer as determined by calibration in a pressurized triaxial chamber
313 Step 2a Foundation soil formation parameter
With the representative value γsoil continue to follow the steps in CPT Material Index through Soil
Behavior Type (SBT) to better define the type of soil at depth 15∙B below Df Use the value of Ic
calculated at depth 15∙B below Df to calculate hs with Equation 36 This parameter is based on the soil
type with typical values shown in Figure 25 The data on silts and sands are considered fully drained
whereas the fissured clay subset may be partially drained to undrained
59
23 ℎ119904 = 28 minus
119868119888 15 32
1+( ) 24
Figure 25 Foundation soil formation parameter hs versus CPT material index Ic (Mayne 2017)
314 Step 3 Soil elastic modulus and Poissonrsquos ratio
Details about the soil such as its elastic modulus (Es) and Poissonrsquos ratio (ν) will be needed for further
calculations A representative value for the Es can be determined from in-situ field tests Values of ν can
be assigned as 02 for drained sands and as 05 for undrained loading cases involving clays (Jardine et al
1985 Burland 1989)
315 Step 4 Net cone tip resistance
Calculate qtnet the mean value of net cone tip resistance 15middotB below the foundation bearing elevation
using Equation 33
33 119902119905119899119890119905 = 119902119905 minus 120590119907119900
60
316 Step 5 Bearing capacity of the soil
Use the assumed B and L calculated hs and qtnet to determine the soils bearing capacity (qmax) from
Equation 34 Use Table 6 to calculate qmax by using the maximum allowable settlement ratio (sB)max
correlating to the soil type If hs is in between the given values interpolate to acquire (sB)max
Table 6 Bearing capacity defined by soil type
Type of Soil hs (sB)max
Clean Sands 058 12
Silts 112 10
Fissured Clays 147 7
Intact Clays 270 4
05 minus0345 119904 119871 119902119898119886119909 = ℎ119904 ∙ 119902119905119899119890119905 ∙ (119861
) ∙ (119861
) 34 119898119886119909
317 Step 6 Settlement
Settlement can be calculated directly using the results from Equation 35 A FS equal to 3 is common in
foundation engineering
2 119902119898119886119909 0345 1 frasl119865119878 119871 119904 = 119861 ∙ [ ∙ ∙ ( ) ] 35
119861 ℎ119904 119902119905119899119890119905
318 Step 7 Final Check
Check that the applied stress (q) is less than qmax using Equation 36 Repeat the process again with a
new B andor new L if q gt qmax
05 119904 ) 119902 = ℎ119904 ∙ 119902119905119899119890119905 ∙ (119861
minus0345 119871 ∙ ( )
11986136
61
32 EXAMPLE PROBLEMS
321 Example 3 Direct CPT Method on Sands
A footing size needs to be determined based on the given CPT data collected for the South Abutment
(Figure 27) The footing stress (q) was determined to be 8000 psf Estimate a footing size (B x L) and
determine the bearing capacity of the foundation (qmax) using the direct CPT method provided Also
determine the expected settlement based on the calculated bearing capacity
Figure 26 Diagram of footing profiles for Example 3
The soil elastic modulus determined from seismic CPT (SCPT) are shown in Table 7
Table 7 SCPT Results
Depth (feet) Es (tsf)
Bottom of layer
3 557
6 433
9 557
12 695
17 590
22 501
27 571
32 505
62
Figure 27 CPT data from Northern Minnesota
Solution
Step 1 Assume the footing is placed below the frost depth of 6 feet
Estimate L L = 50 feet = 600 inches
63
Estimate B B = 12 feet = 144 inches
Estimate footing thickness t t = 2 feet
Df = 6 feet = 72 inches
Df + 15middotB = 24 feet = 288 inches
Step 2 Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 28 Schematic of foundation design associated with CPT data
Layer 1
From Figure 28 fs = 20 psi taken as a representative value of the layer
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
64
Layer 2
From Figure 28 fs = 20 psi taken as a representative value of the layer
20psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
Layer 3
From Figure 28 fs = 8 psi taken as a representative value of the layer
8 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1134 pcf
145 psi
Using the unit weights calculated between Df and 15B determine a representative unit weight of the soil to calculate the total and effective soil stresses Based on layer 3 being the weakest supporting layer γsoil = 113 pcf
σvo = γsoil ∙ (119863119891 + 15 ∙ B) = 113 lbsfraslft3 ∙ 24 feet = 2721 psf = 189 psi
prime σvo = σvo minus uo = 2721 psf minus 0 psf = 2721 psf = 189 psi
Calculate the cone tip resistance between Df and Df + 15B below the foundation depth
qt = qc + u2 ∙ (1 minus a) = 1250 psi + 0 psi ∙ (1 minus 08) = 1000 psi
Step 2a CPT Material Index
Using Figure 28 the sleeve friction at 15B below the foundation depth is about 16 psi
65
fs 16 psi Fr() = 100 ∙ = 100 ∙ = 13
(qt minus σvo) (1250 psi minus 189 psi)
Iterate to solve for Ic Steps are not shown for brevity
(qt minus σvo)σatm (1250 psi minus 189 psi )145 psi = = Qtn prime (σvoσatm)n (189 psi145 psi)n
prime σvo 189 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Qtn = 703 n = 072 lt 10 Ic = 210
66
Figure 29 Soil type for example problem 3
Use the Ic value to determine the foundation soil formation parameter
Calculate the foundation soil formation parameter The tip resistance is about 1250 psi at 24 feet and
the cone area ratio was determined to be 08 from information provided by the manufacturer
23 23 hs = 28 minus = 28 minus = 078 = SandSilt 15 15 Ic 210
1 + ( ) 1 + ( ) 24 24
67
Step 3 Determine representative values of soil elastic modulus from soil testing and Poissons ratio The
soil elastic modulus was taken as the average value between depths of Df and 15middotB (6 feet and 24 feet)
433 + 557 + 695 + 590 + 501 Es = = 555 tsf = 708 psi
5
ldquoDrained sandsiltrdquo gives a ν = 020
Step 4 Calculate the net cone tip resistance
qtnet = qt minus σvo = 1250 psi minus 189 psi = 12311 psi
Step 5 Calculate the bearing capacity of the sand
In drained sandssilts the bearing capacity is taken as the stress when (sB) = 011 (or 11 foundation width)
minus0345 minus0345 05 s L 600 qmax = hs ∙ qtnet ∙ ( ) ∙ ( ) = 078 ∙ (9795) ∙ (011)05 ∙ ( )
B B 144
= 193 psi = 27860 psf qmax
Assuming a factor of safety (FS) of 3
qmax = 645 psi = 9287 psf
3
Step 6 Calculate settlement
68
119902119898119886119909 0345 2
1 frasl119865119878 L 119904 = 119861 ∙ [ ∙ ∙ ( ) ]
ℎ119904 119902119905119899119890119905 B
2 0345 1 645 psi 600 inches s = 144 inches ∙ [ ∙ ∙ ( ) ] = 18 inches
078 12311 psi 144 inches
Step 7 Determine if q gt qmax
q = 8000 psf lt 9287psf = qmax3
69
CHAPTER 4 DIRECT CPT METHOD FOR DEEP FOUNDATIONS
41 INTRODUCTION
The axial compression capacity (Qtotal) for a single pile includes a side component (Qside) end bearing
component (Qbase) and pile weight (Wpile) as shown in Figure 30 Since piles commonly push through
several layers a summation of the unit side frictions acting on the pile segments must be considered
over the length of the pile While the examples shown in this Guide resemble hand calculations
computer software is more efficient Programming the procedure is possible and represents a practical
method of designing deep foundations However commercial software is frequently available and is
MnDOTrsquos most common method of designing deep foundations
There are upwards of 40 different direct CPT methods that have been developed over the past five
decades to determine a piles axial compression capacity Many of the earliest methods relied on hand-
recorded information where qc data from mechanical CPTs would be collected at 20 cm intervals
whereas the direct CPT method uses scaled penetrometer readings via specified algorithms to obtain
the pile unit side friction and unit end bearing The method that will be used for deep foundation design
is the Modified UniCone method which uses all three readings of the electronic piezocone (qt fs and u2)
while addressing a variety of pile foundation types
Figure 30 Direct CPT evaluation of axial pile capacity
70
The modified UniCone Method is based upon a total of 330 pile load tests (three times the original
Unicone database) that were associated with SCPTu data Originally the UniCone method provided
approximate soil classification in five groups via a chart of qE vs fs (Figure 31) where qE = qt -u2 Later
using the modified approach with a larger data set provided soil sub classifications as shown in Figure
32 This new 9-zone normalized soil behavior type is determined using CPT data in combination with the
CPT Material Index
Figure 31 UniCone Method soil behavior type using CPT (Mayne 2017)
Figure 32 Modified UniCone Method soil behavior type using CPT (Mayne 2017)
71
42 MODIFIED UNICONE METHOD
In order to determine the axial pile capacity using the modified method the first step requires the
determination of geoparameters as shown in Direct CPT Method for Soil Characterization Once the
soil unit weight and CPT Material index are determined for each soil layer then the effective cone
resistance pile unit side friction (fp) and pile end bearing resistance (qb) can be determined
421 Step 1
Work through the steps provided in CPT Method for Soil Characterization until each soil layer is
defined by its CPT material index and soil behavior type using Figure 6 After these steps have been
completed continue to Step 2 to determine qE qb and fp
422 Step 2
Once qt is determined for each layer the effective cone resistance can be calculated using Equation 37
119902119864 = 119902119905 minus 1199062 37
Where (a) qE is the specific value at each elevation along the pile sides for determining fp and (b) at the bottom of the pile qE is averaged in the vicinity of the pile tip from the tip bearing elevation to about one diameter beneath the tip for determining qb
Using CPT material index and qE the pile end bearing resistance is calculated using Equation 38
38 119902119887 = 119902119864 ∙ 10(0325∙119868119888minus1218)
The pile unit side friction is obtained from qE and Ic
39 119891119901 = 119902119864 ∙ 120579119875119879 ∙ 120579119879119862 ∙ 120579119877119860119879119864 ∙ 10(0732∙119868119888minus3605)
Where θPT = coefficient for pile type (084 for bored 102 for jacked 113 for driven piles) θTC = coefficient for loading direction (111 for compression and 085 for tension) θRATE = rate coefficient applied to soils in SBT zone 1 through 7 (109 for constant rate of penetration test and 097 for maintained load tests)
72
423 Step 3
Due to piles extending through multiple layers the unit side components acting on various pile
segments would need to be summed if not using the direct CPT method Since CPT calculates data at
regular intervals of 2 cm to 5 cm along the sides of the pile the average fp in each layer can be used
directly in Equation 40 to obtain the shaft capacity
119876119904119894119889119890 = 119891119901 ∙ 119860119904 40
Where fp = average pile side friction along pile length from eqn 39 As = πmiddotdmiddotH d = pile diameter and H = length embedded below grade
The base capacity for a pile in compression loading is given by Equation 41 For piles in tension (or uplift)
Qbase can be taken as 0
= 119902119887 ∙ 119860119887 41 119876119887119886119904119890 Where
qb = end bearing resistance from eqn 38 Ab = πmiddotd24 (area of a circular pile)
424 Step 4
The final step is to calculate the axial pile capacity using Equation 42
119876119905119900119905119886119897 = 119876119904119894119889119890 + 119876119887119886119904119890 minus 119882119901119894119897119890 42
43 AXIAL PILE DISPLACEMENTS
Movement of pile foundations can be assessed using elastic continuum theory which has been
developed using finite element analyses boundary elements and analytical closed-form solutions In
the case of piles passing through several soil layers the elastic solution can be used by stacking pile
segments (each with its own stiffness) as represented by soil Youngrsquos modulus The use of software is
recommended for pile groups Several available programs such as DEFPIG GROUP and PIGLET can
handle pile groups under axial and lateralmoment loading
73
44 EXAMPLE PROBLEMS
441 Example 5 Direct CPT Methods Axial Pile Capacity
Several piles need to be placed beneath the edge of a building Use the CPT data collected for this site
shown in Figure 34 to determine the axial capacity for one of the piles Assume round steel driven piles
will be used with a diameter of 1275 inches wall thickness of 025 inches and lengths of 80 feet
Concrete will be used to fill the piles To solve for all geoparameters use the Direct CPT Method for Soil
Characterization section All sand layers can be assumed ldquodrainedrdquo with ν = 02
Figure 33 Deep foundation end bearing pile diagram
74
Figure 34 CPT data from Minnesota for example problem 5
75
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 35 Soil layers using ldquorules of thumbrdquo for pile capacity example using direct CPT method
Layer 1
From Figure 35 fs = 18 psi taken as a representative value of the layer
76
18 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1210 pcf
145 psi
Layer 2
From Figure 35 fs = 12 psi taken as a representative value of the layer
12psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 3
From Figure 35 fs = 20 psi taken as a representative value of the layer
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
CPT Material Index
Layer 1
From Figure 35 fs = 18 psi and qc = 3000 psi
qt = qc + u2 ∙ (1 minus a) = 3000 psi + 20 psi ∙ (1 minus 08) = 3004 psi
∙ 4 feet = 1219 pcf ∙ 4 feet = 4877 psf = 34 psi σvo = γt
fs 18 psi Fr() = 100 ∙ = 100 ∙ = 060
(qt minus σvo) (3004 psi minus 34 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
77
uo = 0 psi
prime σvo = σvo minus uo = 34 psi minus 0 psi = 34 psi
(qt minus σvo)σatm (3004 psi minus 34 psi )145 psi = = Qtn prime )n (σvoσatm (34 psi145 psi)n
prime σvo 34 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(060)]2
Qtn = 3593 n = 038 lt 10 Ic = 14 (ie sand)
Layer 2
From Figure 35 fs = 12 psi and qc = 500 psi
qt = qc + u2 ∙ (1 minus a) = 500 psi + 40 psi ∙ (1 minus 08) = 508 psi
= 4838 psf + γt ∙ 45 feet = 4838 psf + 1172 pcf ∙ 45 feet = 5756 psf = 400 psi σvo
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 260
(qt minus σvo) (508 psi minus 400 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
78
uo = 0 psi
prime σvo = σvo minus uo = 400 psi minus 0 psi = 400 psi
(qt minus σvo)σatm (508 psi minus 400 psi )145 psi = = Qtn prime (σvoσatm)n (400 psi145 psi)n
prime σvo 400 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(260)]2
Qtn = 117 n = 10 le 10 Ic = 29 (ie clayey silt)
Layer 3
From Figure 35 fs = 20 psi and qc = 5000 psi
qt = qc + u2 ∙ (1 minus a) = 5000 psi + 0 psi ∙ (1 minus 08) = 5000 psi
= 5756 psf + γt ∙ 6 feet = 5756 psf + 1219 pcf ∙ 6 feet = 6488 psf = 451 psi σvo
fs 20 psi Fr() = 100 ∙ = 100 ∙ = 040
(qt minus σvo) (5000 psi minus 451 psi)
79
Step 2 Iterate to solve for Ic Steps are not shown for brevity
prime σvo = σvo minus uo = 451 psi minus 0 psi = 451 psi
(qt minus σvo)σatm (5000 psi minus 451 psi )145 psi = = Qtn prime )n (σvoσatm (451 psi145 psi)n
prime σvo 451 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(040)]2
= 1801 n = 06 lt 10 = 15 (ie sand) Qtn Ic
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic (14) Qtn (359) and Fr (06) the first layer is defined as a ldquoGravelly Sandrdquo
from Figure 36
80
Figure 36 Soil layer 1 using SBT method
81
Layer 2
Based on values of Ic (29) Qtn (117) and Fr (26) the second layer is defined as a ldquoSilty Mixrdquo
from Figure 37
Figure 37 Soil layer 2 using SBT method
82
Layer 3
Based on values of Ic (15) Qtn (180) and Fr (04) the third layer is defined as a ldquoSandrdquo from
Figure 38
Figure 38 Soil layer 3 using SBT method
Effective Cone Resistance
Layer 1
83
qE = qt minus u2 = 3004 psi minus 20 psi = 2984 psi
Layer 2
qE = qt minus u2 = 508 psi minus 40 psi = 468 psi
Layer 3
qE = qt minus u2 = 5000 psi minus 0 psi = 5000 psi
Pile End Bearing Resistance
Layer 3
= qE ∙ 10(0325∙Icminus1218) qb = 5000 psi ∙ 10(0325∙15minus1218) = 9085 psi
Pile Unit Side Friction
Layer 1
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 2984 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙14minus3605) = 99 psi
Layer 2
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 468 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙29minus3605) = 211 psi
Layer 3
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 5000 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙15minus3605) = 202 psi
84
Axial Pile Capacity
The side capacity is based on pile design as shown in Figure 39
Layer 1
Qside = fp ∙ As = 99 psi ∙ (π ∙ d ∙ L) = 99 psi ∙ (π ∙ 1275 in ∙ 48 in) = 19064 lb
Layer 2
Qside = fp ∙ As = 211 psi ∙ (π ∙ d ∙ L) = 211 psi ∙ (π ∙ 1275 in ∙ 588 in) = 457122 lb
Layer 3
Qside = fp ∙ As = 202 psi ∙ (π ∙ d ∙ L) = 202 psi ∙ (π ∙ 1275 in ∙ 240 in) = 58170 lb
Figure 39 Soil layering compared to pilings
85
End Bearing Resistance in Layer 3
d2 1275 in2
Qbase = qb ∙ Ab = 9085 psi ∙ (π ∙ ) = 9085 psi ∙ (π ∙ ) = 115932 lb 4 4
Wp = (γsteel ∙ Asteel ∙ Depth) + (γsteel ∙ Aconc ∙ Depth)
Wp = (490 pcf ∙ 003 ft2 ∙ 55 ft) + (150 pcf ∙ 085 ft2 ∙ 55 ft) = 7724 lbs
Layer 3
Qtotal = ΣQside + Qblayer3 minus Wp
Qtotal = (19064 + 45713 + 58170) lbs + 115932 lbs minus 7724 lbs = 642563 lbs
= 642 kips Qtotal
Table 8 Summary of selected and calculated parameters
Layer L (in) qc
(psi)
fs
(psi)
Fr Ic Qtn qE
(psi)
qb
(psi)
Qbase
(lb)
fp
(psi)
Qside
(lb)
Qtotal
(kip)
1 48 3000 18 06 14 356 2984 NA NA 99 19064
2 588 500 12 26 29 117 468 NA NA 211 457122
3 240 5000 20 04 15 180 5000 9085 115932 202 58170
Total 115932 534355 642
86
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Amar S Baguelin F Canepa Y and Frank R (1998) New design rules for the bearing capacity of shallow foundations based on Menard PMT Geotechnical Site Characterization Vol 2 (Proc ISC-1 Atlanta) 727-733
American Petroleum Institute (API) (1981) Planning designing and constructing fixed offshore platforms Recommended practice API-RP2A 17th edition Washington DC API
American Petroleum Institute (API) (1987) Planning designing and constructing fixed offshore platforms Recommended practice API-RP2A 18th edition Washington DC API
American Petroleum Institute (API) (1989) Planning designing and constructing fixed offshore platforms Recommended practice API-RP2A 19th edition Washington DC API
Andersen KH and Stenhamar P (1982) Static plate loading tests on overconsolidated clay Journal of the Geotechnical Engineering Division ASCE 108 (GT7) 919-934
Basu D and Salgado R (2012) Load and resistance factor design of drilled shafts in sands Journal of Geotechnical amp Geoenvironmental Engineering 138 (12) 1455-1469
Berardi R Jamiolkowski M and Lancellotta R (1991) Settlement of shallow foundations in sands selection of stiffness on the basis of penetration resistance Geotechnical Engineering Congress Vol 1 185-200 (Proc GSP-27 Boulder) Reston Virginia ASCE
Brand EW Muktabhant C and Taechathummarak A (1972) Load tests on small foundations in soft clay Performance of Earth and Earth-Supported Structures Vol 1 Part 2 903-928 (Proc PC) ASCE Reston Virginia ASCE
Briaud JL and Gibbens RM (1999) Behavior of five large spread footings in sand Journal of Geotechnical amp Geoenvironmental Engineering 125 (9) 787-797
Briaud JL (2007) Spread footings in sand load-settlement curve approach Journal of Geotechnical amp Geoenvironmental Engineering 133 (8) 905-920
Brown DA Turner JP and Castelli RJ (2010) Drilled Shafts Construction Procedures and LRFD Design Methods Geotechnical Engineering Circular GEC 10 Report No FHWA NHI-10-016 Washington DC National Highway Institute US DOT
87
Burland J B (1973) Shaft Friction of Piles in Clay -A Simple Fundamental Approach Ground Eng 6(3) 30-42
Cai G Liu S and Puppala A (2012) Reliability assessment of CPTu-based pile capacity predictions in soft clay deposits Engineering Geology 84-91
Canadian Geotechnical Society (1992) Canadian Foundation Engineering Manual Third Edition Vancouver British Columbia BiTech Publishers
Cerato AB and Lutenegger AJ (2007) Scale effects of shallow foundation bearing capacity on granular material Journal of Geotechnical amp Geoenvironmental Engineering 133 (10) 1192-1201
Chen BS-Y and Mayne PW (1996) Statistical relationships between piezocone measurements and stress history of clays Canadian Geotechnical Journal 33 (3) 488-498
Cho W and Finno RJ (2010_ Stress-strain responses of block samples of compressible Chicago glacial clays Journal of Geotechnical amp Geoenvironmental Engineering 136 (1) 178-188
Chung SF Randolph MF and Schneider JA (2006) Effect of penetration rate on penetrometer resistance in clay Journal of Geotechnical amp Geoenvironmental Engineering 132 (9) 1188-1196
Clausen CJF Aas PM and Karlsrud K (2005) Bearing capacity of driven piles in sand the NGI approach Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis Group
Consoli NC Schnaid F and Milititsky J (1998) Interpretation of plate load tests on residual soil site Journal of Geotecnical amp Geoenvironmental Engingeering 124 (9) 857-867
Dasenbrock DD (2006) Assessment of pile capacity by static and dynamic methods to reconcile design predictions with observed performance Proc 54th Minnesota Geotechnical Engineering Conference 95-108 edited by Labuz J Univ of Minnesota St Paul
DeBeer E and Martens A (1957) Method of computation of an upper limit for the influence of heterogeneity of sand layers on the settlement of bridges 275-282 Proc ICSMFE-4 Vol 1 London ICSMFE
DeBeer EE (1965) Bearing capacity and settlement of shallow foundations on sand Proc Symposium on Bearing Capacity and Settlement of Foundations 15-33 Duke University Durham NC
Decourt L (1999) Behavior of foundations under working load conditions Proc PCSMGE-11 Foz do Iguassu Brazil Vol 4 453-487
Demers D and Leroueil S (2002) Evaluation of preconsolidation pressure and the overconsolidation ratio from piezocone tests of clay deposits in Quebec Canadian Geotechnical Journal 39 (1) 174-192
Eslaamizaad S and Robertson PK (1996) Cone penetration test to evaluate bearing capacity of foundations in sand 429-438 Proc CGCFG-49 Vol 1 St Johns Newfoundland
Eslami A Alimirzaei M Aflaki E and Molaabasi H (2017) Deltaic soil behavior classification using CPTu records Marine Georesources amp Geotechnology 35 (1) 62-79
88
Eslami A and Gholami M (2005) Bearing capacity analysis of shallow foundations from CPT data 1463-1466 Proc ICSMGE- 16 Vol 3 (Osaka) Millpress Rotterdam
Eslami A and Gholami M (2006) Analytical model for the ultimate bearing capacity of foundations from cone resistance Scientia Iranica 13 (3) 223-233
Eslami A and Fellenius BH (1997) Pile capacity by direct CPT and CPTu methods applied to 102 case histories Canadian Geotechnical Journal 34 (6) 886-904
Fellenius BH (2009) Basics of Foundation Design Vancouver BiTech Publishers
Fellenius BH and Altaee A (1994) Stress and settlement of footings in sand Vertical and Horizontal Deformations of Foundations and Embankments 2 1760-1773
Fellenius BH and Eslami A (2000) Soil profile interpreted from CPTu data Proc Geotechnical Engineering Conference Year 2000 Geotechnics Asian Inst of Technology Bangkok Thailand Available from wwwfelleniusnet
Finno R J and Chung C K (1992) Stress-strain-strength responses of compressible Chicago glacial clays Journal of Geotechnical Engineering 118 (10) 1607ndash1625
Fleming K Weltman A Randolph M and Elson K (2009) Piling Engineering Third Edition London Taylor amp Francis Group
Frank R and Magnan J-P (1995) Cone penetration testing in France national report Proceedings International Symposium on Cone Penetration Testing 3 (CPT95 Linkoumlping) Swedish Geotechnical Society Report 3(95) 147-156
Gavin K Adekunte A and OKelly B (2009) A field investigation of vertical footing response on sand Geotechnical Engineering 162 257-267
Hegazy YA (1998) Delineating geostratigraphy by cluster analysis of piezocone data PhD dissertation School of Civil amp Environmental Engineering Georgia Institute of Technology Atlanta GA
Holtz RD Kovacs WD and Sheehan TC (2011) An Introduction to Geotechnical Engineering 2nd Edition London Pearson Publishing
Hunt CE Pestana JM Bray JD and Riemer M (2002) Effect of pile driving on static and dynamic properties of soft clay Journal of Geotechnical amp Geoenvironmental Engineering 128 (1) 13-24
Jamiolkowski M Ladd CC Germaine JT and Lancellotta R (1985) New developments in field and lab testing of soils Proc ICSMFE-11 1 57-154
Jardine R Chow F Overy R and Standing J (2005) ICP design methods for driven piles in sands and clays London Thomas Telford Publishing
Jardine RJ Lehane BM Smith P and Gildea P (1995) Vertical loading experiments on rigid pad foundations at Bothkennar Geotechnique 45 (4) 573-597
89
Karlsrud K (2012) Prediction of load-displacement behavior and capacity of axially-loaded piles in clay based on analyses and interpretation of pile load test results PhD Dissertation Norwegian Univ Science amp Technology Trondheim Norway
Karlsrud K Clausen CJF and Aas PM (2005) Bearing capacity of driven piles in clay the NGI approach Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis
Karlsrud K Lunne T Kort DA and Strandvik S (2001) CPTU correlations for clays Proc 16th ICSMGE 2 683-702
Kimura T Kusakabe O and Saitoh K (1985) Geotechnical model tests of bearing capacity problems in a centrifuge Geotechnique 35 (1) 33-45
Knappett JA and Craig RF (2012) Craigs Soil Mechanics 8th Edition London Spon Press Taylor amp Francis
Kolk HJ and Velde EVD (1996) A reliable method to determine friction capacity of piles driven into clays Paper No 7993 Proc Offshore Technology Conference Houston
Kolk HJ Baaijens AE and Senders M (2005) Design criteria for pipe piles in silica sands Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis Group
Kulhawy FH (2004) On the axial behavior of drilled shafts Geosupport 2004 1-18
Kulhawy FH and Jackson CS (1989) Foundation Engineering Current Principles amp Practices 2 (GSP 22) Reston VA ASCE
Kulhawy FH and Mayne PW (1990) Manual for Estimating Soil Properties for Foundation Design EPRI Report EL-6800 Electric Power Research Institute Palo Alto 306 page Download from wwwepricom
Kulhawy FH Trautmann CH Beech JF ORourke TD and McGuire W (1983) Transmission Line Structure Foundations for Uplift-Compression Loading Report EL-2870 Palo Alto CA Electric Power Research Institute wwwepricom
Ku T and Mayne PW (2015) In-situ lateral stress coefficient (K0) from shear wave velocity measurements in soils Journal of Geotechnical amp Geoenvironmental Engineering 141 (12) 101061(ASCE) GT1943-56060001354
Ladd CC and DeGroot DJ (2003) Recommended practice for soft ground site characterization Soil amp Rock America 2003 3-57
Larsson R (1997) Investigations and Load Tests in Silty Soils SGI Report R-54 Linkoumlping Swedish Geotechnical Institute
Larsson R (2001) Investigations and Load Tests in Clay Till SGI Report R-59 Linkoumlping Swedish Geotechnical Institute
Lee J and Salgado R (2005) Estimation of bearing capacity of circular footings on sands based on cone penetration tests Journal of Geotechnical amp Geoenvironmental Engineering 131 (4) 442-452
90
Lehane BM (2003) Vertically loaded shallow foundation on soft clayey silt Geotechnical Engineering 156 (1) Thomas Telford London 17-26
Lehane BM (2013) Relating foundation capacity in sands to CPT qc Geotechnical amp Geophysical Site Characterization 1 London Taylor amp Francis Group
Lehane BM Doherty JP and Schneider JA (2008) Settlement prediction for footings on sand Deformational Characteristics of Geomaterials (IS-Atlanta) 1 Rotterdam Millpress
Lehane BM Li Y and Williams R (2013) Shaft capacity of displacement piles in clay using the CPT Journal of Geotechnical amp Geoenvironmental Engineering 139 (2) 253-266
Lehane BM Schneider JA and Xu X (2005) The UWA-05 method for prediction of axial capacity of driven piles in sand Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis
Low HE Lunne T Andersen KH Sjursen MA Li X and Randolph MF 2010 Estimation of intact and remoulded undrained shear strengths from penetration tests in soft clays Geotechnique 60 (11) 843-859
Lunne T Robertson PK and Powell JJM (1997) Cone Penetration Testing in Geotechnical Practice London RoutledgeBlackie Academic
Lutenegger AJ and DeGroot DJ (1995) Settlement of Shallow Foundations on Granular Soils Report Contract 6332 Task Order 4 submitted to Massachusetts Highway Dept by University of Mass-Amherst Amherst MA
Lutenegger AJ and Adams MT (2003) Characteristic load-settlement behavior of shallow foundations Proc FONDSUP 2 381-392
Mase T and Hashiguchi K (2009) Numerical analysis of footing settlement problem by subloading surface model Soils amp Foundations 49 (2) 207-220
Mayne PW (1987) Determining preconsolidation stress and penetration pore pressures from DMT contact pressures Geotechnical Testing Journal 10(3) 146-150
Mayne PW (1991) Determination of OCR in Clays by Piezocone Tests Using Cavity Expansion and Critical State Concepts Soils and Foundations 31 (2) 65-76
Mayne PW (2007) NCHRP Synthesis 368 on Cone Penetration Test Washington DC Transportation Research Board National Academies Press wwwtrborg
Mayne PW (2009) Geoengineering Design Using the Cone Penetration Test Richmond BC Canada ConeTec Inc
Mayne PW (2014) Keynote Interpretation of geotechnical parameters from seismic piezocone tests Proceedings 3rd Intl Symp on Cone Penetration Testing 47-73
Mayne PW (2016) Evaluating effective stress parameters and undrained shear strengths of soft-firm clays from CPT and DMT Australian Geomechanics Journal 51 (4) 27-55
91
Mayne PW (2017) Stress history of soils from cone penetration tests 34th Manual Rocha Lecture Soils amp Rocks Vol 40 (3) Saouml Paulo ABMS
Mayne PW Christopher B Berg R and DeJong J (2002) Subsurface Investigations -Geotechnical Site Characterization Publication No FHWA-NHI-01-031 Washington DC National Highway Institute Federal Highway Administration
Mayne PW Coop MR Springman S Huang A-B and Zornberg J (2009) Geomaterial Behavior and Testing Proc 17th Intl Conf Soil Mechanics amp Geotechnical Engineering 4 Rotterdam MillpressIOS Press
Mayne PW and Illingworth F (2010) Direct CPT method for footings on sand using a database approach Proc ISCPT-2 3 315-322
Mayne PW and Niazi F (2017) Recent developments and applications in field investigations for deep foundations Proceedings 3rd Bolivian Conference on Deep Foundations 1 141-160
Mayne PW Peuchen J and Baltoukas D (2015) Piezocone evaluation of undrained strength in soft to firm offshore clays Frontiers in Offshore Geotechnics III 2 1091-1096
Mayne PW and Poulos HG (1999) Approximate displacement influence factors for elastic shallow foundation systems ASCE Journal of Geotechnical amp Geoenvironmental Engineering 125 (6) 453-460
Mayne PW Uzielli M and Illingworth F (2012) Shallow footing response on sands using a direct method based on cone penetration tests Full Scale Testing and Foundation Design (GSP 227) Reston Virginia ASCE
Mayne PW and Woeller DJ (2014) Direct CPT method for footings on sands silts and clays Geocharacterization for Modeling and Sustainability (GSP 234) 1983-1997
McClelland B (1974) Design of Deep Penetration Piles for Ocean Structures Journal Geot Eng Div 100(GT7) 705-747
Mesri G (1975) Discussion on new design procedure for stability of soft clays ASCE Journal of the Geotechnical Engineering Division 101(GT4) pp 409-412
Meyerhof GG (1956) Penetration tests and bearing capacity of cohesionless soils Journal of Soil Mechanics amp Foundations Division 82 (SM1) 1-19
Meyerhof GG (1974) General Report Penetration testing outside Europe Proceedings ESOPT-1 Vol 21 Swedish Geotechnical Society Stockholm
Meyerhof GG (1976) Bearing capacity and settlement of pile foundations Journal of Geotechnical Engineering Division 102(3) 197-228
Missiaen T Verhegge J Heirman K and Crobe P (2015) Potential of cone penetrating testing for mapping deeply buried palaeolandscapes in the context of archaeological surveys in polder areas Journal of Archaeological Science 55 174-187
92
Mlynarek Z Wierzbicki J Goglik S and Bogucki M (2014) Shear strength and deformation parameters of peat and gyttja from CPTu DMT and VST Proceedings 5th International Workshop on CPTu and DMT in soft clays and organic soils 193-210 Poznan Poland Polish Committee on Geotechnics Menard Polska Tensar Internationa
Nejaim PF Jannuzzi GMF and Danziger FAB (2016) Soil behavior type of the Sarapuiacute II test site Geotechnical amp Geophysical Site Characterization 5 1009-1014
Niazi FS and Mayne PW (2013) Cone penetration test based direct methods for evaluating static axial capacity of single piles Geotechnical and Geological Engineering 31 (4) 979-1009
Niazi FS and Mayne PW (2015) Enhanced Unicone expressions for axial pile capacity evaluation from piezocone tests Proc International Foundations Conference amp Equipment Expo RestonVA ASCE
Niazi FS and Mayne PW (2016) CPTu-based enhanced UniCone method for pile capacity Engineering Geology 212 21-34
Nowacki F Karlsrud K and Sparrevik P (1996) Comparison of recent tests on OC clay and implications for design Large Scale Pile Tests in Clay London Thomas Telford
ONeill MW (2001) Side resistance in piles and drilled shafts Journal of Geotechnical amp Geoenvironmental Engineering 127 (1) 3-16
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Pestana JM Hunt CE and Bray JD (2002) Soil deformation and excess pore pressure filed around a closed-ended pile Journal of Geotechnical amp Geoenvironmental Engineering 128 (1) 1-12
Poulos HG and Davis EH (1974) Elastic Solutions for Soil and Rock Mechanics New York Wiley amp Sons
Poulos HG and Davis E (1980) Pile Foundation Analysis amp Design New York John Wiley amp Sons
Powell JJM and Lunne T (2005) Use of CPTU data in clays and fine-grained soils Studia Geotechnica et Mechanica XXVII(3) 29-66
Randolph MF (2003) Rankine Lecture Science and empiricism in pile foundation design Geotechnique 53 (10) 847-875
Robertson PK (1990 1991) Soil classification using the cone penetration test Canadian Geotechnical Journal 27 (1) 151-158 Closure 28 (1) 176-178
Robertson PK (1991) Estimation of foundation settlements in sand from CPT Geotechnical Engineering Congress 2 Reston Virginia ASCE
93
Robertson PK (2009) Cone penetration testing a unified approach Canadian Geotechnical J 46 (11) 1337-1355
Robertson PK (2009a) Performance based earthquake design using the CPT Keynote Lecture International Conference on Performance-Based Design in Earthquake Geotechnical Engineering IS Tokyo Tsukuba Japan
Robertson PK (2009b) Interpretation of cone penetration tests a unified approach Canadian Geotechnical Journal 46 (11) 1337-1355
Robertson PK and Cabal KL (2007) Guide to cone penetration testing Second Edition Signal Hill California Gregg In-Situ
Robertson PK and Cabal KL (2015) Guide to Cone Penetration Testing for Geotechnical Engineering 6th Edition Signal Hill California Gregg Drilling amp Testing Inc
Schmertmann JH (1970) Static cone to compute static settlement over sand Journal of the Soil Mechanics and Foundations Division 95 (SM3) 1011-1043
Schmertmann JH (1978) Guidelines for cone penetration test performance and design Report FHWA-TS-78-209 Washington DC Federal Highway Administration
Schnaid F Wood WR Smith AKC and Jubb P (1993) Investigation of bearing capacity and settlements of soft clay deposits at Shellhaven Predictive Soil Mechanics London Thomas Telford
Schneider JA Xu X and Lehane BM (2008) Database assessment of CPT-based design methods for axial capacity of driven piles in siliceous sands J Geotechnical and Geoenvironmental Engineering 134 (9) 1227-1244
Semple D (1980) Recent Developments in the Design amp Construction of Piles London Institution of Civil Engineers
Senneset K Sandven R amp Janbu N (1989) Evaluation of soil parameters from piezocone tests In-Situ Testing of Soil Properties for Transportation Transportation Research Record 1235 24-37
Shahri AA Melehmir A and Juhlin C (2015) Soil classification analysis based on piezocone penetration test data Engineering Geology 189 32-47
Stuedlein AW and Holtz RD (2010) Undrained displacement behavior of spread footings in clay The Art of Foundation Engineering Practice GSP 198 Reston Virginia ASCE
Takesue K Sasao H and Matsumoto T (1998) Correlation between ultimate pile skin friction and CPT data Geotechnical Site Characterization 2 1177ndash1182
Tand KE Funegard EG and Briaud J-L (1986) Bearing capacity of footings on clay CPT method Use of In-Situ Tests in Geotechnical Engineering GSP 6 1017-1033
Tand KE Warden PE and Funegaringrd EG (1995) Predicted-measured bearing capacity of shallow footings on sand PISCPT 2 589-594
Thomas D (1968) Deep soundings test results and the settlement of spread footings on normally consolidated sands Geotechnique 18 (2) 472-488
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Tomlinson MJ (1957) The adhesion of piles driven into clay soils Proc 4th Intl Conf on Soil Mechanics and Foundation Engineering 2 66-71
Tomlinson MJ (1986) Foundation Design amp Construction London Longman Scientific
Tomlinson MJ (1987) Pile Design and Construction Practice London Viewpoint Publication
Trofimenkov JG (1974) Penetration testing in the USSR Proceedings ESOPT-1 1 147-154
Tumay MT Hatipkarasulu Y Marx ER and Cotton B (2013) CPTPCPT-based organic material profiling Proc 18th Intl Conf on Soil Mechanics amp Geotechnical Engineering Paris 633-636
Uzielli M and Mayne PW (2011) Statistical characterization and stochastic simulation of load-displacement behavior of shallow footings Proc Georisk 2011 Geotechnical Risk Management amp Assessment (GSP 224) 672-679
Uzielli M and Mayne PW (2012) Load-displacement uncertainty of vertically-loaded shallow footings on sands and effects on probabilistic settlement estimation Georisk Assessment and Management of Risk for Engineered Systems and GeoHazards 6 (1) 50-69
Valsson SM (2016) Detecting quick clay with CPTu Proceedings 17th Nordic Geotechnical Meeting Reykajavik Iceland Challenges in Nordic Geotechnics
Van Dijk BFJ and Kolk HJ (2011) CPT-based design method for axial capacity of offshore piles in clay Frontiers in Offshore Geotechnics II 555-560
Vesić AS (1975) Bearing capacity of shallow foundations Foundation Engineering Handbook New York Van Nostrand Reinhold Publishers
Vesic AS (1977) NCHRP Synthesis of Highway Practice 42 Design of Pile Foundations Washington DC Transportation Research Board
Yu F and Yang J (2012) Base capacity of open-ended steel pipe piles in sand J Geotechnical and Geoenvironmental Engineering 138 (9) 1116-1128
Zawrzykraj P Rydelek P and Bakowska A (2017) Geoengineering properties of Eemian peats from Radsymin (central Poland) in the light of static cone penetration and dilatometer tests Engineering Geology 226 290-300
95
APPENDIX A
List of Figures
Figure A1 Conventional methods versus direct CPT approach to shallow foundation response A-3
Figure A2 Direct bearing capacity relationship for embedded square and strip footings situated on clean
sands (after Schmertmann 1978) A-6
Figure A3 Direct CPT bearing factor Rk as function of footing width B and embedment depth from finite
element analyses (Tand et al 1995) A-6
Figure A4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio and sand
consistency (Eslaamizaad amp Robertson 1996) A-7
Figure A5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp Salgado (2005) in
terms of footing size relative density and base settlement A-8
Figure A6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and normalized
cone tip resistance (after Eslami amp Gholami 2005) A-8
Figure A7 Direct relationship between ultimate bearing stress in clays and measured cone tip resistance
(Schmertmann 1978) A-10
Figure A8 Direct relationship between ultimate foundation bearing stress in clays and net cone tip
resistance (after Tand et al 1986) A-11
Figure A9 Measured load-displacements for each of the five large footings at TAMU (Briaud amp Gibbens
1994) sand site A-12
Figure A10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU (Briaud amp
Gibbens 1994) footings A-13
Figure A11 Characteristic stress versus square root normalized displacements for TAMU footings A-15
Figure A12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sand A-16
Figure A13 Summary of sand formation factors with corresponding CPT resistances A-19
Figure A14 Normalized foundation stresses vs square root of normalized displacement for spread
footings on sand (modified after Mayne et al 2012) A-20
Figure A15 Definitions of capacity from three types of observed foundation load test responses
(modified after Kulhawy 2004) A-21
Figure A16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footings A-22
A-1
Figure A17 Measured load vs displacement curves for 3 shallow foundations on clay at Baytown Texas
(Stuedlein amp Holtz 2010) A-25
Figure A18 Characteristic stress vs square root of normalized displacement for all three foundations at
Baytown site (Stuedlein amp Holtz 2010) A-26
Figure A19 Normalized foundation stress vs square root of normalized displacements A-27
Figure A20 Applied foundation stress vs CPT-calculated stress for 67 footings A-28
Figure A21 Applied footing stress vs CPT calculated stress on logarithmic scales A-29
Figure A22 Summary graph and equations of direct CPT method for evaluating the vertical A-30
Figure A23 Foundation soil formation parameter hs versus CPT material index Ic A-32
Figure A24 Bearing capacity ratio qmaxqnet versus CPT material index Ic A-33
Figure A25 Influence factor for rectangular foundations from elastic theory solution (Mayne amp
Dasenbrock 2017) A-34
List of Tables
Table A1 Direct CPT methods for bearing capacity of footings on clean sands A-4
Table A2 Direct CPT methods for bearing capacity of footings on clays A-9
Table A3 Summary of large footings soil conditions and reference sources of database A-17
Table A4 Sources for shallow foundation performace A-34
A-2
A1 INTERPRETATION OF SOIL PARAMETERS FROM CONE
PENETRATION TESTS
A11 Introduction
Geotechnical site characterization is an important first step towards the evaluation of
subsurface conditions and determination of soil layering geomaterial classification and the
evaluation of soil engineering parameters for the analysis and design of foundations retaining
walls tunnels excavations embankments and slope stability Increasing use of the electronic
cone penetrometer in highway applications is prominent in the USA because the results are
obtained much faster and less expensive than traditional methods that rely on rotary drilling
augering sampling and laboratory testing Moreover the cone penetration test (CPT) provides
at least three independent and continuous readings on soil behavior that are digitally recorded
and fully available at the end of the sounding As such the reliable and consistent
interpretation of CPT data is important since civil engineering works will best realize the
efficiency and economy of this technology in constructed facilities for county state and
interstate projects
A111 Cone Penetration Testing
The cone penetrometer is an electronically-instrumented steel probe that is vertically pushed into the
ground at constant rate of 20 mms (08 insec) using a hydraulic system The penetrometer is outfitted
with load cells pressure transducers and inclinometers that measure at least three readings
continuously with depth (a) cone tip resistance qt (b) sleeve friction fs and (c) porewater pressure u2
With the latter reading the device is called a piezocone thus the designation CPTu is used to indicate
porewater pressure Additional sensors are available that can be used to measure inclination
resistivity pH temperature shear wave velocity and other readings if desired
Figure A1 shows a schematic rendering of the cone penetration test that is performed in the
field per ASTM D 5778 procedures The hydraulic system is nominally rated at 180 kN (20 tons)
capacity and often mounted on a truck but can also be positioned on tracked vehicles or
anchored frames Figure A2 shows a MnDOT cone truck that uses the full dead weight of the
rig for the hydraulic reaction forces which are applied at the center of the vehicle The cab is
enclosed so that soundings may proceed during inclement weather and the data acquisition
system and operator are protected
Figure A 1 Setup and procedures for co ne penetration testing (CPT)
A-4
Figure A2 CPT truck-mounted rig used by MnDOT
A representative sounding taken at the I-35 test site just northeast of Saint Paul is presented in Figure
A3 Here the separate profiles of qt fs and u2 with depth are shown in side-by-side plots In the last
graph the hydrostatic porewater pressure (uo) due to the groundwater table is indicated by the blue
dashed line
These readings are used to interpret the soil layers types and properties of the ground as
discussed in the following sections
A-5
Figure A3 Representative CPT sounding at I-35 test site northeast of St Paul MN showing profiles of
(a) cone resistance (b) sleeve friction and (c) penetration porewater pressures
A112 Soil Parameter Interpretation
A variety of soil engineering parameters can be interpreted from CPT results on the basis of
theoretical frameworks analytical models or numerical simulations otherwise by empirical
methods based on correlations and statistics of databases Figure A4 shows a conceptual
utilization of the readings from the CPT for interpretation of geostratigraphy compressibility
flow characteristics and stress-strain-strength behavior of soils
Table A1 lists a selection of geoparameters that have been addressed for these purposes The
most common or useful relationships will be discussed in subsequent sections and reference is
made to other sources (Lunne et al 1997 Mayne 2007 Robertson amp Cabal 2015) for additional
details
A-6
Figure A4 Conceptual post-processing of CPT results for geoparameter evaluation
Table A1 List of common geoparameters interpreted from CPT data
Symbol Parameter Remarks Notes
SBT Soil behavior type (SBT)
Ic CPT material index
γt Unit weight
ρt Mass Density
σvo Overburden stress
u0 Hydrostatic pressure
A-7
σvo Effective stress
Table A1 Continued
Symbol Parameter Remarks Notes
σp Preconsolidation stress σp = 033 (qnet)m where m = fctn (Ic)
(or Effective Yield Stress) where stresses are in kPa
YSR Yield stress ratio YSR = σpσvo (formerly OCR)
c Effective cohesion intercept
ϕ Effective friction angle
su Undrained shear strength
D Constrained modulus
G0 Small-strain modulus
G Shear modulus
E Youngs modulus
cvh Coefficient of consolidation
K Bulk modulus
A-8
LI
K0 Lateral stress coefficient
k Hydraulic conductivity
Liquefaction index
A12 Geostratigraphy and Soil Behavioral Type (SBT)
In routine CPT soil samples are not normally taken and thus the measured stress friction andor
porewater pressure readings are used to infer the types of soils This can be accomplished using
three basic approaches
a Rules of Thumb or approximate guidelines for a quick visual assessment
b Soil Behavioral Type (SBT) Charts that are based on either the three readings (qt fs u2) net readings
including net cone resistance (qnet = (qt-σvo) effective cone resistance (qE = qt - u2) friction ratio (Rf =
100middotfsqt) and excess porewater pressures or using normalized CPT readings such as Q F Bq or U
c Probabilistic methods where the CPT readings have been calibrated from lab tests on
recovered soil samples (eg Tumay et al 2013)
A121 Approximate Rules of Thumb
The approximate rules of thumb provide simple guidelines for soil type and suggest that sands are
identified when qt gt 5 MPa and u2 asymp u0 whereas intact clays are prevalent when qt lt 5 MPa and u2 gt u0
(Mayne et al 2002) The magnitude of porewater pressures reflects the consistency of the intact clay
such that soft (u2 asymp 2middotu0) firm (u2 asymp 4middotu0) stiff (u2 asymp 8middotu0) and hard (u2 asymp 20middotu0) However for fissured
overconsolidated clays the measured porewater pressures are less than hydrostatic in fact often
negative u2 lt 0 On land negative porewater pressure readings at the shoulder filter element of the
piezocone should be greater than -1 bar ie u2 gt -100 kPa (-147 psi) Also for clean sands the friction
ratio (FR = 100middotfsqt lt 1) while for insensitive clays (FR gt 4)
A-9
Figure A5 presents a 36-m deep piezocone sounding from the MnDOT Wakota Bridge site which crosses
the Mississippi River at I-494 (Dasenbrock 2006) The qt and u2 readings have been annotated using the
aforementioned rules of thumb indicating the predominance of sands at this site As noted there are
five interbedded clay layers evident at depths of 1 m 3-4 m 7 - 12 m 17 m and 22-27 m
Figure A5 Interpretation of soil types from CPT data taken from the Wakota Bridge on I-494 using
approximate rules of thumb
A122 Soil Behavioral Type Charts
The most common approach for identification of soil types is based on soil behavioral charts
Reviews of several chart methods are provided elsewhere (Kulhawy amp Mayne 1990 Fellenius amp
Eslami 2000 Shahri et al 2015) Current popularity favors the 9-zone classification system to
evaluate soil behavioral type (SBT) that uses normalized piezocone readings (Robertson 1990
1991 Lunne et al 1997)
A-10
(119902119905 minus 120590119907119900) (a) normalized tip resistance 119876 = A11 prime 120590119907119900
100∙119891119904 (b) normalized sleeve friction 119865() = A12
(119902119905 minus 120590119907119900)
(1199062 minus 1199060) (c) normalized porewater pressure 119861119902 = A13
(119902119905 minus 120590119907119900)
While the Q-F-Bq classification is a three-dimensional plot it is often presented as a set of two-dimensional
plots specifically Q vs F and Q vs Bq as shown in Figure A6 Data are grouped according to layers and
can be superimposed to identify their association with the nine soil behavioral types All indirect CPT soil
classification approaches should be cross-checked and verified for a particular geologic setting and local
geotechnical conditions before routine use in practice
Figure A6 Nine-zone soil behavioral type charts (a) normalized cone resistance vs normalized
sleeve friction (b) normalized cone resistance vs normalized porewater pressure parameters
(adapted after Robertson 1991 Lunne et al 1997)
The development of a CPT material index (Ic) has been found advantageous in the initial screening of soil
types and helps to organize the sounding into their respective SBT zones of similar soil response In this
case the CPT index is found from (Robertson 2009)
A-11
A2 22 )log221()log473( rtnc FQI
where Qtn = modfied stress-normalized net cone resistance given by
(119902119905 minus 120590119907119900)120590119886119905119898 119876 = A31 prime (120590119907119900120590119886119905119898)119899
where σatm = reference stress equal to atmospheric pressure (1 atm asymp 1 bar = 100 kPa asymp 1 tsf asymp 147 psi)
The exponent n is soil-type dependent n = 1 (clays) n asymp 075 (silts) and n asymp 05 (sands) If units of bars are
used then Qtn (units of bars) is simply
(119902119905 minus 120590119907119900) 119876 = 119899 A32
prime ) (120590119907119900
The operational value of exponent n is found by iteration Initially an exponent n = 1 is used to calculate
the starting value of Ic (ie Qtn = Q) and then the exponent is upgraded to
prime 120590119907119900 119899 = 0381 ∙ 119868119888 + 005 ( ) minus 015 le 10 A4 120590119886119905119898
Then the index Ic is recalculated Iteration converges quickly and generally only 3 cycles are needed to
secure the operational Ic at each depth The soil zones and associated Ic values are detailed in Figure A7
The sensitive soils of zone 1 can be screened using the following expression
A-12
Zone 1 119876119905119899 lt 12119890119909119901( minus14 ∙ 119865119903)
A5
If any soil layers are found within zone 1 (sensitive soils) then caution should be exercised as these
structured clays are prone to instability collapses and difficulties in construction and performance
Another indicator of zone 1 sensitive soils is when the normalized porewater parameter Bq gt 08 When
these criteria are evident the geoengineer should consult with senior geotechnical personnel or the chief
engineer for guidance
Figure A7 Delineation of soil behavioral type zones using CPT material index Ic
A-13
Stiff overconsolidated clayey sands and sandy clays soils of zone 8 (15 lt Fr lt 45) and zone 9 (Fr gt
45) can be identified from the following criterion
Zones 8 and 9 0020)1(00030)1(0050
12
rr
tnFF
Q A6
The remaining soil types are identified by the CPT material index Zone 2 (organic clayey soils Ic ge 360)
Zone 3 (clays to silty clays 295 le Ic lt 360) Zone 4 (silt mixtures 260 le Ic lt 295) Zone 5 (sand mixtures
205 le Ic lt 260) Zone 6 (clean sands 131 le Ic lt 205) and Zone 7 (gravelly to dense sands Ic le 131)
The red dashed line at Ic = 260 represents an approximate boundary separating drained (Ic lt 260) from
undrained behavior (Ic gt 260)
Once the specific zone has been assigned to a layer a visual representation can be made to show
either the zone number or colorization so that the predominant layers and soil types can be
realized
A123 Probabilistic Soil Types from CPT
The inference of soil type has been addressed using probabilistic methods as discussed by Abu-
Farshakh et al (2008) and Tumay et al (2013) Here the CPT readings can be post-processed to
provide the percentage of sand silt and clay components with depth The output is consistent
and compatible with the current MnDOT soil classification chart (Figure A8) which is similar in
content to that used by the USDA
A-14
Figure A8 Chart for soil classification system by MnDOT
A13 Identifying Soft Organic Soils
The identification of soil type using cone penetration tests (CPT) is usually done on the basis of empirical
soil behavioral type (SBT) charts that use the measured readings (qt fs andor u2) While there at least
25 different sets of charts available (ie Hegazy 1998 Eslami et al 2017 Valsson 2016) one of the most
widely used and popular is the 9-zone SBT system that uses three normalized cone readings Qtn Fr and
Bq (Robertson amp Cabal 2007 Lunne et al 1997) The system is denoted as SBTn and plotted on charts in
terms of Q vs F and Q vs Bq as shown in Figure A9
A-15
1
10
100
1000
01 1 10
Norm
aliz
ed
Con
e T
ip R
esis
tan
ce
Q
Normalized Friction Fr = 100 fst(qt - svo) ()
9-Zone Soil Behavioral Type Chart
Gravelly Sands(zone 7)
Sands
(zone 6)
Sandy Mixtures(zone 5)
Silt Mix(zone 4)
Clays
(zone 3)
Organic Soils(zone 2)
Sensitive Clays
and Silts(zone 1)
Very stiff OC clayto silt (zone 9)
Very stiffOC sandto clayey
sand(zone 8)
1
10
100
1000
-06 -04 -02 00 02 04 06 08 10 12 14
No
rmal
ized
Co
ne
Tip
Res
ista
nce
Q
Normalized Porewater Bq = Du2(qt-svo)
Clays(zone 3)
Sensitive(zone 1)Organic
(zone 2)
Gravelly Sands(zone 7)
Sands(zone 6)
Figure A9 CPT charts for SBTn system (a) Q versus F (b) Q versus Bq
Of particular concern in geotechnical site characterization is the proper identification of soft
organic soils such as organic clays and silts (OL and OH) and peats (Pt) These soils often cause
difficulties and problems during construction and performance of civil engineering works because
of bio-degradation long-term creep excessive settlements andor other issues
The SBTn system has a category labelled Zone 2 recognizes organic soils However several recent
studies have noted that data from CPTu soundings do not always fall within the bounds
delineated using Zone 2 even though laboratory tests and field inspections clearly note the
presence of organic soils Lab tests to classify organic soils includes loss on ignition (LOI gt 5)
and two pairs of Atterberg limits testings per ASTM standards as well as the visual-manual
identification of strong organic odor and smell and dark color notably black and dark gray
Results by Mlynarek et al (2014) show that organic soils in Poland are not adequately identified by the
Q-F chart as seen in Figure A10 Moreover studies by Zawrzykraj et al (2017) found that neither the Q-
A-16
F and Q-Bq plots worked well to recognize organic soils and peats in Poland Nejaim et al (2016) found
that Brazilian soft organic clays were misclassified as Zone 3 (clays) rather than Zone 2 (organic clays)
Similarly a recent study on CPTu data from organic clays located at 24 sites in the USA Sweden Mexico
Brazil and Australia have been compiled (Agaiby 2018) These data are shown to generally avoid falling
with the Zone 2 boundaries in either Q-F or Q-Bq charts thus are not recognized during these soundings
as indicated by Figure A11 The exception in this case is the Mexico City clay which is correctly identified
however the other 23 sites are generally not properly recognized and categorized Additional findings by
Missiaen et al (2015) found that the CPTu results in Belgian soils also did not identify organic clays and
peats in the non-normalized version of these charts that uses Qt versus friction ratio (Rf = 100middotfsqt) as
detailed by Robertson amp Cabal (2007)
Figure A10 Soil behavioral chart with superimposed CPTu data from Polish organic soils
(from Mlynarek et al 2014)
A-17
Figure A11 Paired SBTn charts with CPTu data from 24 organic clay sites indicating non-compliance
with Zone 2 (organic soils)
(compiled by Agaiby 2018)
A14 Simplified Yield Stress Evaluation in Clays by CPTu
From the development of a hybrid formulation of spherical cavity expansion (SCE) theory with critical-
state soil mechanics (CSSM) the effective preconsolidation stress or yield stress (σp) of clays can be
expressed in terms of three piezocone parameters (a) net cone tip resistance qnet = qt - σvo (b)
measured excess porewater pressure Δu2 = u2 - u0 and (c) effective cone resistance qE = qt - u2 Details
on the SCE-CSSM solutions are given elsewhere (Mayne 1991 Mayne 2017 Agaiby amp Mayne 2018)
A-18
For normal and well-behaved clays which include inorganic fine-grained soils such as clays and silts
of low sensitivity a set of simple relationships can be derived By adopting characteristic default values
for effective friction angle ϕ = 30deg and rigidity index (IR = 100) linear expressions for the effective
preconsolidation stress are obtained
prime 120590119901 = A7 033 ∙ 119902119899119890119905
prime 120590119901 = 053 ∙ ∆1199062 A8
prime 120590119901 = 060 ∙ 119902119864 A9
A141 Application to soft Chicago clay
An example of a common geomaterial in this category includes the infamous soft Chicago clays
that were deposited in a glacio-lacustrine environment (Cho amp Finno 2010) The soft clay has a
mean water content of 20 liquid limit of 38 and plasticity index PI= 12 Results of CPTu tests
conducted at the national geotechnical experimentation site (NGES) at Northwestern University
are shown in Figure A12 (Ouyang amp Mayne 2017) As seen in the profile a thin soft clay resides
between depths of 9 to 105 m with a thicker soft clay layer in the range of 12 to 18 m
A-19
0
2
4
6
8
10
12
14
16
18
20
00 02 04 06 08 10 12D
ep
th (
m)
CPTu qt and u2 (kPa)
sand (SP)
Blodgett
Park Ridge
Deerfieldsoft clay
ClayLayerof Study
Northwestern UnivNational Test Site
Figure A12 Results of CPTu sounding at the NGES at Northwestern University Illinois
Post-processing of the CPTu data using Equations A7 A8 and A9 show good agreement in evaluating the
profile of effective yield stress in this clay layer compared with results from one-dimensional
consolidation tests made on undisturbed samples taken at the site
A-20
10
11
12
13
14
15
16
17
18
19
20
0 100 200 300 400 500 600 700 800 900 1000
De
pth
(m
)Effective Yield Stress sp (kPa)
Consols
033 qnet
053 Du2
060 qE
svo
Figure A13 Evaluation of effective yield stresses at NWU using consolidation tests and CPTu
A142 Application to soft Bay Mud California
Another example of a well-behaved fine-grained soil is the San Francisco Bay Mud which is a soft clay
deposit in northern California Results of a representative CPTu are shown in Figure A14 and indicate the
soil profile at a site in eastern San Francisco (Hunt et al 2002) For this site index values include 70 lt
wn lt 90 60 lt LL lt 80 and 35 lt PI lt 45 Post-processing the CPTu data using Equations A7 A8 and
A9 show good agreement with each other as well as with the results of consolidation tests as seen in
A-21
Figure A15 The consolidation tests were performed using a CRS device as reported by Pestana et al
(2002)
San Francisco Bay Mud (Pestana et al 2002)
0
5
10
15
20
25
30
35
40
0 1000 2000 3000
De
pth
(m
ete
rs)
Cone Resistance qt (kPa)
0
5
10
15
20
25
30
35
40
0 10 20 30 40 50 60
Sleeve Friction fs (kPa)
0
5
10
15
20
25
30
35
40
0 500 1000 1500
Porewater u2 (kPa)
Miscellaneous Fill
Young Bay Mud
Young Bay Mud
Clayey Sand
Miscellaneous Fill
Young Bay Mud
Young Bay Mud
Figure A14 Representative CPTu in San Francisco Bay Mud (data from Hunt et al 2002)
A-22
San Francisco Bay Mud (Pestana et al 2002)
0
5
10
15
20
25
30
0 100 200 300 400 500 600
De
pth
(m
ete
rs)
Preconsolidation Stress sp (kPa)
svo
033 qnet
053 Du2
060 qE
CRS
svo
Figure A15 Evaluation of effective yield stresses in Bay Mud from consolidation tests and CPTu
A15 CPTu Screening of Soft Organic Clays
For soft organic clays the Equations A7 A8 and A8 will not apply thus can serve as a means to
screen such geomaterials from the soil profile Two examples of CPTu soundings in soft organic
clays are presented to illustrate the approach Additional case records and documentation of
CPTu data in organic clays are given in Agaiby (2018)
A151 Soft organic alluvial soils in Washington DC
Results of in-situ tests including CPTu soundings in soft organic clayey silts along the Potomac River at the
Anacostia Naval Air Station are presented by Mayne (1987) Here the soft soils are categorized as organic
clayey silts (OH) per the Unified Soils Classification Systems (USCS) with mean values (and plus and minus
A-23
one standard deviations) of wn = 68 plusmn 16 LL = 83 plusmn 25 and PI = 37 plusmn 17 A representative
piezocone sounding is depicted in Figure A16 For the post-processing of CPTu data the use of the Q-F
and Q-Bq graphs do not find any points within the Zone 2 category for organic soils When the data are
evaluated to determine the profile of effective yield stress the Equations A7 A8 and A9 do not agree as
shown by Figure A17
0
5
10
15
20
25
0 2000 4000 6000
Dep
th (
met
ers)
Cone Resistance qt (kPa)
0
5
10
15
20
25
0 20 40 60 80 100
Sleeve Friction fs (kPa)
0
5
10
15
20
25
0 200 400 600
Pore Pressure u2 (kPa)
Figure A16 Representative CPTu in soft organic clay along the Potomac River DC
A-24
0
5
10
15
20
25
0 3 6 9D
epth
(m
eter
s)Soil Behaviour Type Zone
Q and Fchart
0
5
10
15
20
25
0 100 200 300 400 500 600
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
033qnet
053 Du2
06 qE
Oed
Figure A17 Post-processing of CPTu data from Anacostia-Bolling (a) SBTn zones using CPT material
index Ic (b) CPTu screening equations for yield stress
Note that the hierarchy of the results from Equations A7 A8 and A9 show that the CPTu
evaluations for preconsolidation stress in soft organic clays indicates
[A10] σp = 053 Δu2 lt 033 qnet lt 060 qE
A152 Soft organic peaty clay in Saint Paul MN
Results of CPTu tests were collected during a training exercise underneath I-35E just northeast
of Saint Paul MN in 2007 All three MnDOT CPT rigs available at the time were used to obtain
in-situ test data at the site The site is well known as having soft organic soils and samples were
obtained from three borings made at the site (Boring ID T523 T542 and T518) Boring logs
A-25
indicate the presence of highly organic silt loams with marl and spongy organic clayey silts
From 12 samples the natural water contents ranged from 30 to 191 with a mean of 112 plusmn
58
A representative piezocone sounding from the site (MnDOT CPTu ID F22Y0703C) is used for
this illustration and is presented as Figure A18 Post-processing these data using the SBTn
algorithms and CPT material index show that the soils classify primarily as Zone 3 (clays) and
Zone 4 (silty mix) thus do not identify the geomaterials properly as Zone 2 (organic soils)
0
5
10
15
0 500 1000 1500 2000
Dep
th (
met
ers)
Cone Tip Resistance
qt (kPa)
0
5
10
15
0 10 20 30 40 50 60
Sleeve Friction
fs (kPa)
0
5
10
15
-100 0 100 200 300 400
Porewater Pressure
u2 (kPa)
Figure A18 MnDOT CPTu sounding at the I-35E test site near St Paul Minnesota
A-26
1
10
100
1000
01 1 10
Norm
aliz
ed
Con
e T
ip R
esis
tan
ce
Q
Normalized Friction Fr = 100 fst(qt - svo) ()
9-Zone Soil Behavioral Type Chart
St PaulGravelly Sands(zone 7)
Sands
(zone 6)
Sandy Mixtures(zone 5)
Silt Mix(zone 4)
Clays
(zone 3)
Organic Soils(zone 2)
Sensitive Clays
and Silts(zone 1)
Very stiff OC clayto silt (zone 9)
Very stiffOC sandto clayey
sand(zone 8)
1
10
100
1000
-06 -04 -02 00 02 04 06 08 10 12 14
No
rmal
ized
Co
ne
Tip
Res
ista
nce
Q
Normalized Porewater Bq = Du2(qt-svo)
Clays(zone 3)
Sensitive(zone 1)Organic
(zone 2)
Gravelly Sands(zone 7)
Sands(zone 6)
Figure A19 Post-processing St Paul sounding using the 9-zone SBTn classification system
Using the recommended approach Figure A20 shows the CPT-evaluated yield stress profiles from
equations A7 A8 and A9 Here again the three estimates of σp do not agree but show the same
hierarchy as detailed in Equation A10
A-27
0
5
10
15
20
0 50 100 150 200 250 300 350 400
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
06 qeff svo
033qnet Consols
054 Du2
Figure A20 CPTu screening of soft organic soils at St Paul test site
A16 CPTu Evaluations of Yield Stress in Soft Organic Soils
Once the presence of soft organic clays and silts is identified using the aforementioned CPT
screening algorithms the evaluation of effective yield stress is conducted using the procedure
outlined in Chapter 2 of this MnDOT manual For soft organic soils an exponent of m = 090 is
recommended (Mayne et al 2009) Therefore the preconsolidation stress is obtained from
(units of kPa)
prime = A11 120590119901 033 ∙ (119902119899119890119905)090
The algorithm can be expressed in dimensionless form by
prime = A12 120590119901 033 ∙ (119902119905 minus 120590119907119900)119898prime ∙ (120590119886119905119898100)1minus119898prime
A-28
where σatm = reference stress (= 100 kPa = 147 psi) and m = 090 for soft organic silts and clays
The approach is applied to the two former example case studies in Figure A21 (Washington DC)
and Figure A22 (St Paul MN) respectively showing good agreement with benchmark results
taken from laboratory consolidation tests on undisturbed samples of these sites in both cases
0
5
10
15
20
25
0 100 200 300 400 500 600
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
033qnet^090
Oed
Figure A21 Profiles of yield stress from consolidation tests and CPTu method for organic clayey silts at
Anacostia-Bolling site in Washington DC
A-29
0
5
10
15
20
0 20 40 60 80 100 120 140 160
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
CPT
Consols
120590119901prime = 033 119902119899119890119905
090 in kPa
Figure A22 Profiles of yield stress from consolidation tests and CPTu method for organic clayey silts at
MnDOT test site near Saint Paul Minnesota
A17 Soil Unit Weight
As shown by Figure A23 total soil unit weight (t) can be estimated from the CPT sleeve friction resistance
(Mayne 2014) The equation can be expressed by
120574119905 = 120574119908 ∙ [122 + 015 ∙ 119897119899 (100 lowast 119891119904120590119886119905119898 + 001)] A13
where w = unit weight of water and satm = atmospheric pressure
A-30
Figure A23 Evaluation of soil unit weight from CPT sleeve friction
A18 Effective Stress Friction Angle
Sands
The effective stress friction angle () is one of the most important soil properties as it governs the strength
of geomaterials as well as affects soil-pile interface and pile side friction While an effective cohesion
intercept (c) can also be considered this is usually reserved for cemented or bonded geomaterials or
unsaturated soils and may lose its magnitude with time ageing or with prolonged environmental
changes
For clean quartz to silica sands where porewater pressures are essentially hydrostatic (Bq = 0) the
following expression has been calibrated with triaxial compression test results from undisturbed sand
samples and normalized cone resistances as presented in Figure A24 (Mayne 2007 2014)
A-31
120601prime(119889119890119892) = 176deg + 110deg 119897119900119892 (1199021199051) A14
Figure A24 Evaluation of effective stress friction angle in quartz-silica sands from CPT
Note Relationship applies to drained soil behavior when Ic lt 26 andor Bq lt 01
where qt1 is an earlier form of stress normalized cone tip resistance for sands given by
(119902119905 120590119886119905119898) 1199021199051 = A151 05 prime (120590119907119900120590119886119905119898)
A-32
In terms of units of bars this is expressed simply as (in units of bars)
119902119905 1199021199051 = 05 A152 prime ) (120590119907119900
Recently Robertson amp Cabal (2015) recommended the use of the modified form of normalized cone
resistance (Qtn) in Equation A10
120601prime (119889119890119892) = 176deg + 110deg 119897119900119892 (119876119905119899) A16
In sands qnet asymp qt since the overburden term is small relative to the cone tip resistance Figure A25 shows
that the two normalizations give very comparable results
Figure A25 Comparable values from qt1 and Qtn normalization schemes for CPTs in sands
Clays and Silts
A-33
In the case of soft to firm intact clays and silty clays the effective friction angles are determined from the
normalized cone resistance and porewater pressure parameters (Senneset et al 1989 Mayne 2016) as
shown in Figure A26 The exact solution when the angle of plastification = 0 is given as
qBQ
)tan1(tan61
1)tanexp()245(tan 2
A17
Approximate NTH Solution for from CPTu
qBQ
)tan1(tan61
1)tanexp()245(tan2
Approx asymp 295deg∙Bq0121∙[0256 + 0336∙Bq + log Q ]
1
10
100
20 25 30 35 40 45
Co
ne
Res
ista
nce
Nu
mb
er Q
(degrees)
Bq
010204060810
c = 0 = 0deg
Theory (dots)
Approximation (lines)
Theory
Figure A26 Evaluation of effective stress friction angle in clays and silts from CPTu results
Note Relationship applies to undrained soil behavior when Ic ge 26 andor Bq ge 01
which can be approximately inverted into the form (Mayne 2007)
A-34
0121 120601prime = 295deg ∙ 119861119902 ∙ [ 0256 + 0336 ∙ 119861119902 + log(119876)] A18
This algorithm is specifically applicable for the following ranges of porewater pressure parameter (01 le
Bq lt 1) and effective stress friction angles (20deg le lt 45deg)
A19 Stress History
The stress history can be characterized by an apparent yield stress or preconsolidation stress (sp) as well
as by its normalized and dimensionless form YSR = spsvo = yield stress ratio The YSR is in effect the
same as overconsolidation ratio (OCR) however now generalized to accommodate mechanisms of
preconsolidation that occur beyond just erosion glaciation and removal of overburden stresses but also
due to ageing desiccation repeated cycles of wetting-drying bonding repeated freeze-thaw cycles
groundwater changes and other factors
A-35
A generalized approach for evaluating the yield stress or preconsolidation in soils using net cone
resistance has been formulated as presented in Figure A27 (Mayne 2015) Yield stress in soils from CPT
10
100
1000
10000
10 100 1000 10000 100000
Yie
ld S
tre
ss
s
y
(kP
a)
Net Cone Resistance qt - svo (kPa)
Blessington
General Trend
sy = 033(qt-svo)m
Fissured Clays m = 11
Intact clays m = 10 Sensitive clays m = 09
Silt Mixtures m = 085Silty Sands m = 080
Clean Sands m = 072
m = 11 10 09
085
080
072
Units kPa
Fissured Clays
Figure A27 Evaluation of yield stress or preconsolidation stress in soils from CPT
The algorithm can be expressed in dimensionless form by
1minus119898prime prime 120590119886119905119898 120590119901 = 033 ∙ (119902119905 minus 120590119907119900)119898prime
( ) A19 100
where m = exponent depends on soil type m = 1 (intact clays) 085 (silts) 080 (sandy silts to silty sands)
and 072 (sands) For fissured clays the exponent m may be 11 or higher depending upon the age of the
A-36
formation and degree of jointing and discontinuities Note that fissured clays can be identified when Ic lt
26 and Bq lt 01
The exponent m has also been calibrated with CPT material index as presented in Figure A28
An algorithm to express this relationship for non-fissured soils is given by
25)652(1
2801
cIm
A20
This allows the post-processing of CPT to automatically choose the appropriate exponent m for
a layer by layer analysis
Figure A28 Yield stress exponent m in terms of CPT material index Ic
A-37
A110 Lateral Stress Coefficient
The horizontal geostatic state of stress is represented by the lateral stress coefficient K0 = shosvo
commonly referred to as the at-rest condition The magnitude of K0 for soils that have been loaded and
unloaded can be approximately estimated from
119870119900 = (1 minus 119904119894119899120601prime) ∙ 119874119862119877119904119894119899120601prime A21
Data compiled from in-situ K0 measurements using self-boring pressuremeter tests (SBPMT) total stress
cells (TSC) or push-in spade cells andor laboratory methods (instrumented consolidometers triaxials
andor suction measurements) on a variety of clays silts and sands have been compiled and reported by
Ku amp Mayne (2015) as shown in Figure A29 verifying Equation A15 as a means for evaluating K0 in soils
Figure A29 Relationship between lateral stress coefficient K0 and YSR or OCR in soils (Ku amp Mayne
2015)
A-38
A111 Undrained Shear Strength
The loading of soils can occur under conditions of being fully drained (Du = 0) partially drained or full
undrained (DVV0) where Du = excess porewater pressures (above hydrostatic) and DVV0 = volumetric
strain Further details on specific stress paths are best explained in terms of critical state soil mechanics
or CSSM (Mayne et al 2009 Holtz Kovacs amp Sheahan 2011) The prevailing drainage conditions depend
upon the rate of loading and permeability characteristics of the soil Normally in sands that are pervious
and exhibit high permeability a drained response occurs Exception to this may occur in loose sands
during fast earthquake loading resulting in soil liquefaction In clays that exhibit low permeability a fast
rate of loading will result in undrained loading at constant volume This in fact is a temporary and
transient condition often termed short term loading In the long term eventually porewater pressures
will dissipate (albeit slowly) and a drained response will prevail thus termed long term loading
The overall soil strength is controlled by the effective stress strength envelope Most commonly this is
represented by a simple linear relationship termed the Mohr-Coulomb criterion where the maximum
shear stress (max called the shear strength) is given as
= 119888prime + 120590prime ∙ 119905119886119899 120601prime A22 120591119898119886119909
where c = effective cohesion intercept s = effective normal stress and = effective stress friction angle
As a starting point values of c = 0 and = 30deg can be adopted for all soil types at least until the specific
soil type geologic formation and results from high-quality laboratory or field test data are available
Peak Undrained Shear Strength from Stress History
Using simplified critical state soil mechanics the peak undrained shear strength (max = su) can be
evaluated from the effective stress strength envelope (c = 0 effective friction angle ) and stress history
(ie OCR) in the form
prime DSS 119904119906 = (119904119894119899120601prime2) ∙ (119874119862119877)Λ ∙ 120590119907119900 A23
A-39
which corresponds to a simple shear mode Figure A30 presents the aforementioned relationship
together with data from 17 different clays tested in simple shear
Figure A30 Normalized undrained shear strength from simple shear tests on various clays
versus YSR or OCR
For the triaxial compression (TC) mode the equation would give slightly higher strengths that are
calculated from
prime TC 119904119906 = (1198721198882) ∙ (1198741198621198772)Λ ∙ 120590119907119900 A24
where Mc = (6 middot sin)(3-sin)
Peak Undrained Shear Strength from CPTu
A more direct approach to assessing undrained shear strength is via bearing capacity theory
whereby
A-40
A25 kt
votu
N
qs
s
where Nkt = bearing factor that depends on mode of shearing sensitivity of the clay degree of
overconsolidation and other factors For soft offshore clays the back figured Nkt from data collected at
14 well-documented sites (Low et al 2010) determined mean values based on mode of shearing Nkt =
119 (triaxial compression) Nkt = 136 (simple shear) and Nkt = 133 (vane)
A recent study of 51 clays that were tested by both field piezocone and laboratory CAUC triaxial tests
showed that essentially Nkt = 12 for intact clays and clayey silts of low to medium sensitivity (Mayne
Peuchen amp Baltoukas 2015) Figure A31 shows the slightly different trends for offshore versus onshore
clays For sensitive clays a lower value of Nkt = 10 would be appropriate and for fissured clays a higher
value of Nkt would be in the range of 20 to 30
A-41
A-42
Figure A31 Database from 51 clays with CPT qnet versus lab measured triaxial shear strength
(Mayne Peuchen amp Baltoukas 2015)
For soft to firm clays an alternate means to evaluate undrained shear strength is via the excess porewater
pressures ( u = u2 - u0) from the following
u
uN
us
D
D 2
A26
where N u = porewater bearing factor also dependent on the aforementioned factors The study by
Low et al (2010) found N u = 59 (triaxial compression) N u = 69 (simple shear) and N u = 71 (vane
shear)
A third alternate is found from the effective cone resistance (qE = qt - u2) whereby
kE
tu
N
uqs 2
A27
where NkE is a bearing term for this approach Mayne et al (2015) indicated a mean value of NkE = 8 for
soft-firm intact clays
Remolded Undrained Shear Strength from CPT
The remolded undrained shear strength (sur) is obtained from either field vane tests lab mini-vane shear
tests or lab fall cone devices This affords the evaluation of the clay sensitivity (St) which is defined as the
ratio of peak to remolded strengths at a given water content
119904119906(119901119890119886119896) 119878119905 = frasl A28 119904119906119903
For the CPT the sleeve friction has been noted to give values that are comparable to the
remolded undrained shear strength (Powell amp Lunne 2015) Therefore
asymp 119891119904 A29 119904119906119903
A112 Ground Stiffness and Soil Moduli
The stiffness of the ground can be represented by several geoparameters including the compressibility
indices (Cr Cc Cs) spring constants (kv) and moduli Regarding the latter there are several moduli that
are used in geotechnical engineering including shear modulus (G and Gu) Youngs modulus (E and Eu)
constrained modulus (D) bulk modulus (K) resilient modulus (MR) and subgrade reaction modulus (ks)
A-43
Soil Modulus
The definition of modulus is taken as E = DsD with Figure A32 showing a typical deviator stress (q = s1
- s3) versus axial strain (a) curve Theoretical interrelationships between the elastic moduli G E D and
K have a dependence on the Poissons ratio v The value of Poissons ratio can be taken as vu = 05 for
undrained loading (ie constant volume) while for drained loading which is accompanied by volumetric
strains a value of v = 02 may be used If we adopt the reference modulus as E then the
interrelationships with the other elastic moduli are given by
a Reference Stiffness 119864prime = drained Youngs modulus A30
119864prime
b Shear Modulus 119866prime = A31 [2(1+120584prime)]
119864prime (1minus120584prime) c Constrained Modulus 119863prime = A32
[(1+120584prime)(1minus2120584prime)]
119864prime
d Bulk Modulus 119870prime = A33 [3middot(1minus2120584prime)]
A-44
Figure A32 Representative stress-strain-strength curve for soils in triaxial compression
For the extreme case when = 0 in fact D = E When = 02 the two moduli are only 10
different D = 11middotE thus the constrained modulus and drained Youngs modulus are often
considered somewhat interchangeable
The resilient modulus (MR) applies to pavement analysis and design most commonly measured by cyclic
triaxial testing under repeated load applications In fact MR is a special case of Youngs modulus that
relates to the small strain stiffness measured in the nondestructive range but has developed permanent
plastic strains after many cycles of loading (Brown 1996 Dehler amp Labuz 2007)
The subgrade reaction modulus is actually a combined soil-structural parameter as its value
depends on the ground stiffness and the size of the loaded element The subgrade modulus is
defined as ks = q where q = applied stress and = measured deflection In terms of elasticity
solutions the deflection of a flexible circle of diameter d is given by = qmiddotdmiddot(1-2)E Therefore
119864prime
119896119904 = A34 [119889middot(1minus1205842)]
which has units of kNm3 or pcf
Table A2 summarizes several selected studies towards the approximate evaluation of these moduli
obtained directly from measured CPT data Note however that the results from in-situ penetrometer
tests generally represent a peak strength as indicated in Figure A32 Thus the measured cone tip
resistance (qt) reflects the top of the stress-strain curve either the undrained strength in clays (su) or the
effective friction angle () of sands or even a strength intermediate between these two values Thus
A-45
Source Soils Studied Reference Modulus
Findings
Kulhawy amp Mayne 1990
8 stiff OC clays Lab constrained modulus
D asymp 825 middot (qt - svo)
Mohammed et al (2000)
Resilient modulus
Abu-Farsakh 2004
7 Louisiana sites
Lab constrained modulus
D asymp 358 middot (qt - svo)
Mayne (2007b)
All soil types Constrained moduli from lab consolidation
First-order estimate D asymp middot(qt - svo)
Robertson (2009)
All soil types Constrained modulus from consolidation
D = m middot (qt - svo)
when Ic gt 22
1 m = Qtn when Qtn le 14
2 m = 14 when Qtn gt 14
when Ic lt 22 then
= 003 middot 10055 Ic + 168 m
Liu et al (2016)
16 clay sites in China
Resilient modulus MR = (146qt 053 + 1355 fs
14 + 236)244
(all in MPa)
Casey et al (2016)
8 clays Triaxial tests for undrained Youngs modulus
Eu(svc)07 = 316 - 23 middot LL
where Eu and svc (MPa)
and LL = liquid limit ()
qt is probably not the best means to obtain the slope of the stress-strain curve Instead the SCPTu that
provides the initial tangent shear modulus (Gmax) from the shear wave velocity (Vs) is a better choice
Table A2 Selected Modulus Relationships with Cone Penetration Test Measurements
A-46
Nonlinear Modulus
The small-strain shear modulus (Gmax or G0) represents the initial stiffness of all soils and rocks It is the
beginning of all stress-strain-strength curves for geomaterials and obtained from elasticity theory
119866119898119886119909 = 1198660 = 120588119905 middot 1198811199042 A35
where Vs = shear wave velocity t = tga = total soil mass density t = total soil unit weight and ga =
gravitational acceleration constant (98 ms2)
From a general viewpoint on stiffness the shear modulus G of soil is defined as the slope of shear stress
versus shear strain G = DDs for a tangent definition and by G = s as a secant definition The shear
modulus is related to its associated Youngs modulus E = 2G(1+v) Both moduli are in fact highly
nonlinear ranging from a maximum value at the small-strain stiffness (Gmax = G0 = t middotVs2 where t =
total soil mass density and Vs = shear wave velocity) to intermediate G values at medium strains (asymp 1)
to low values at peak strength As such a variety of algorithms and formulae have been developed to
represent either a partial range or the full stress-strain-strength behavior of soils over a range of
interests (eg Mayne 2005) In these formulations a variable number of input parameters may be
required in order to produce a stress-strain-strength curve
A modified hyperbolic form suggested by Fahey amp Carter (1993) has favorable attributes in that the
modulus reduction factor can be established with only a single variable This allows the initial stiffness
to the be small-strain stiffness (G0) that is reduced in terms of level of mobilized strength eg max =
qqmax = 1FS which is simply the reciprocal of the calculated factor of safety (FS) The magnitude of
secant shear modulus (G) corresponding to the particular level of loading is given by
119866 = 119872119877119865 middot 1198660 = (1198661198660) middot 1198660 A36
A-47
where the MRF = modulus reduction factor determined as
1 minus (120591 )119892 119872119877119865 = (1198661198660) = frasl A37 120591119898119886119909
and g = empirical exponent term Figure A33 shows the relationships for normalized shear stress ()
versus shear strain (s) and corresponding normalized secant shear modulus (G = s) with shear strain
over a range of exponent values 01 le g le 10 For a discussion of tangent G please see Fahey amp Carter
(1993)
Using laboratory data from resonant column-torsional shear tests and triaxial specimens with local
strain measurements for Gmax reference values modulus reduction curves for a selection of sands and
clays are presented in Figure A34 The data indicate that the exponent g falls within the range 02 lt g lt
05 for many soils tested drained and undrained A mean value of g = 03 is recommended for
preliminary site investigations and designs until additional information can be obtained
The value of max is the shear strength of the soil generally taken as either (1) drained (c = 0) max =
svo middot tan or (2) undrained where max = su = undrained shear strength as discussed earlier Thus each
shear stress () can be associated with its shear modulus (G) and the relevant shear strain is found from
s = G This allows for generation of nonlinear stress-strain-strength curves at all depths from SCPTu
data in clays sands and mixed soil types
A-48
Modified Hyperbola (Fahey amp Carter 1993 Canadian Geotech J) Coefficient Term f = 1
``
0
01
02
03
04
05
06
07
08
09
1
000001 00001 0001 001 01 1
No
rmal
ized
GG
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
0
01
02
03
04
05
06
07
08
09
1
0 01 02 03 04 05 06 07 08 09 1
No
rmali
zed
GG
ma
x
Mobilized Shear Stress max
g =1
07
05
03
02
01
0
01
02
03
04
05
06
07
08
09
1
00001 0001 001 01 1
No
rmali
zed
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
0
01
02
03
04
05
06
07
08
09
1
0 1 2 3 4 5 6 7 8 9 10
No
rmali
zed
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
Reduction Factor
GGmax = 1- (max)g
Figure A33 Normalized modulus (GGmax) and normalized shear stress (max) versus shear strain (s)
for modified hyperbolic relationship of Fahey amp Carter (1993)
A-49
0
01
02
03
04
05
06
07
08
09
1
0 01 02 03 04 05 06 07 08 09 1
Mo
du
lus
Re
du
cti
on
G
Gm
ax
or
EE
ma
x
Mobilized Strength max or qqmax
NC SLB Sand
OC SLB Sand
Hamaoka Sand (e=0625)
Hamaoka Sand
Toyoura Sand e = 067
Toyoura Sand e = 083
Ham River Sand
Ticino Sand
Kentucky Clayey Sand
Kaolin
Fujinomori Clay
Pietrafitta Clay
London Clay
Vallericca Clay
Pisa Clay
Pisa Clay
Pisa Clay
Pisa Clay
Kiyohoro Silty Clay
Thanet Clay
Exp g = 02
Exp g = 03
Exp g = 04
Exp g = 05
MRF = 1 - (qqmax)g
Figure A34 Modulus Reduction Factors (MRF) with mobilized strength of clays and sands
A113 Coefficient of Consolidation
The rate at which foundation and embankment settlements occur as well as the dissipation of excess
porewater pressures is controlled by the coefficient of consolidation (cv) The magnitude of cv is also
required in the design of vertical wick drains that can be installed in soft ground to expedite the time for
consolidation Using the results of CPTu dissipation tests which measure the rate at which the u2 readings
vary with time the in-situ profile of cv can be evaluated Often the piezocone-interpreted values of cv
are validated by comparison with results from laboratory one-dimensional consolidation tests (eg Abu-
Farsakh MY 2004) Alternatively a better method is to cross-check the values with the measured full-
scale performance of instrumented embankments that are constructed over the ground and the recorded
time-rate of consolidation and settlements can provide the best cv for that site (Abu-Farsakh MY et al
2011)
A-50
Monotonic Dissipation Tests
An illustrative dissipation curve from piezocone testing at IDT State Highway 95 at an embankment and
bridge crossing is presented in Figure A35 After the CPTu sounding was halted at a depth of 512 feet
the recorded decay of porewater pressures was observed to be monotonic with time until the dissipations
were ended at 1000 seconds During that time the u2 readings decreased from 115 psi to 47 psi At this
location the depth to the groundwater table is about 16 feet thus the calculated equilibrium water
pressure is 15 psi Commonly a characteristic time for piezo-dissipations is taken at 50 degree of
consolidation although other degrees may be adopted By evaluating the value of u2 at 50 as shown in
Figure A35 the characteristic t50 = 404 s can be obtained
0
20
40
60
80
100
120
1 10 100 1000 10000
Pore
wat
er P
ress
ure
u2
(psi
)
Time (s)
CPTU-02E-1 z = 1560 m
uo (hydrostatic)
u2 (initial) = 115 psi
u2 (final projected) = u0 = 15 psi
u2 at 50 = frac12(115+15) = 65 psi
Idaho DOT State Route 95Dissipation at 512 feet
t50 = 404 seconds
Time to reach50 consolidation
A-51
Figure A35 Piezo-dissipation at Sandpoint Bridge Crossing Idaho for depth z = 512 feet illustrating the
evaluation of characteristic time t50
It is also common to plot normalized excess porewater pressures relative to the measured initial Du2
that is obtained during penetration at the constant rate of 20 mms ie DuDui versus time Figure A36
shows a set of piezo-dissipation records for a bridge and embankment site in southern Louisiana
reported by the LTRC While the porewater pressure axis is arithmetic the time scale is often plotted on
either logarithmic or square root scales In either case the t50 is then found from the measured
dissipation curve when DuDui = 050
A-52
Figure A36 Set of monotonic piezo-dissipations taken at various depths for Courtableau Bridge
Site on Louisana State Highway 103 (Abu-Farsakh et al 2011)
Dilatory Dissipation Curves
In a number of situations a monotonic type dissipation does not occur on the onset but instead a dilatory
type curve is observed Here after stopping the push of the penetrometer the recorded porewater
pressures initially begin to rise and eventually reach a peak value then afterwards follow a phase where
the Du values decrease to hydrostatic (Figure A37) A full solution is available for both cases (Burns amp
Mayne 2002) Dilatory responses are most often associated with overconsolidated clays and silts
although occasionally are observed in fine-grained soils with low OCRs
The selection of the characteristic t50 value may found by using a square root of time plot to find the initial
u2 value by projection of the post-peak dissipatory data back to the ordinate axis as shown by the example
in Figure A38 In this example the initial u2 = 480 kPa and then the same procedures in Figure A35may
apply
A-53
Figure A37 Monotonic and dilatory porewater dissipation responses in soils
(after Burns amp Mayne 1998)
Figure A38 Method for obtaining t50 from dilatory dissipation curves
(Schneider amp Hotstream 2010)
A-54
Interpretation of Dissipation Test Data
There are a number of different solutions available for the interpretation of the coefficient of
consolidation (cv) from the piezocone dissipation curves For monotonic type response the strain path
method (SPM) developed at Oxford University (Teh amp Houlsby 1991) is well-recognized while the Georgia
Tech solution by Burns amp Mayne (2002) is based on spherical cavity expansion and critical state soil
mechanics (SCE-CSSM) and handles both monotonic and dilatory curves For both solutions a simplified
procedure can be recommended that relies on the aforementioned t50 value obtained from the measured
field data as given in Table A3 The units for cv are in cm2s as shown or equivalent
Table A3 Recommended procedures for calculating cv from t50 obtained in dissipation tests
Method Equation for coefficient of
consolidation cv
RemarksNotes Eqn
No
SPM
(Teh amp
Houlsby
1991) 50
2)(2450
t
Iac
Rc
v
t50 = measured time to reach 50
dissipation
IR = Gsu = undrained rigidity index
G = shear modulus
su = undrained shear strength
ac = penetrometer radius (ac = 178
cm for 10-cm2 cone ac = 220 for a
15-cm2 size)
A32
SCE-CSSM
(Burns amp
Mayne 2002) 50
7502 )()(0300
t
Iac Rc
v
A33
A-55
Evaluating rigidity index
Both the SPM and SCE-CSSM solutions require an estimate of the in-situ rigidity index (IR = Gsu) of the
soil If the results of SCPTU are available Krage et al (2014) have derived an expression for IR that
depends upon the small-strain shear modulus net cone tip resistance and effective overburden stress
250750
max50
8111
vonet
Rq
GI
s A38
where consistent units are input for Gmax qnet and svo The value of Gmax is obtained from B29 using the
measured shear wave velocity (Vs) and unit weight evaluated from B7
If the shear wave velocity is not measured it can be estimated from the CPT data A number of
methods have been reviewed for CALTRANS by Wair et al (2012) including one that is applicable
to sands silts and clays of low sensitivity that are inorganic and uncemented
119881 (119898119904) = [101 middot 119897119900119892(119902119905) minus 114]167 middot (100 middot 119891119904119902119905)03 A39 119904
where qt is in units of kPa (Hegazy amp Mayne 1995) An alternative relationship is given by Robertson amp
Cabal (2015)
A114 Hydraulic Conductivity
The hydraulic conductivity is a parameter that expresses the flow characteristics of the soil In
geotechnics it is also called the coefficient of permeability (k) and has units of cms or feetday
Through consolidation theory the hydraulic conductivity relates directly to the coefficient of
consolidation
A-56
A40 D
ck wv
where w = unit weight of water and D = constrained modulus
Therefore one approach to evaluating k can be from the site-specific cv obtained from the
dissipation tests using an estimate of D from one of the relationships given in Table A2
An approach developed for soft normally-consolidated soils that uses t50 directly to assess the magnitude
of k is presented in
Figure A39 (Parez and Fauriel 1988) The mean trendline through the wedges of soil type can be
approximated by
251
50 (sec)251
1)(
tscmkh
A41
A-57
Figure A39 Hydraulic conductivity versus dissipation time for 50 consolidation
(Parez amp Fauriel 1988)
A-58
APPENDIX B
List of Figures
Figure B1 Conventional methods vs direct CPT approach to shallow foundation response B-3
Figure B2 Direct bearing capacity relationship for embedded square and strip footings situated on clean
sands (after Schmertmann 1978) B-6
Figure B3 Direct CPT bearing factor Rk as function of footing width B and embedment depth from finite
element analyses (Tand et al 1995) B-6
Figure B4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio and sand
consistency (Eslaamizaad amp Robertson 1996) B-7
Figure B5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp Salgado (2005) in
terms of footing size relative density and base settlement B-8
Figure B6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and normalized
cone tip resistance (after Eslami amp Gholami 2005) B-8
Figure B7 Direct relationship between ultimate bearing stress in clays and measured cone tip resistance
(Schmertmann 1978) B-10
Figure B8 Direct relationship between ultimate foundation bearing stress in clays and net cone tip
resistance (after Tand et al 1986) B-11
Figure B9 Measured load-displacements for each of the five large footings at TAMU (Briaud amp Gibbens
1994) sand site B-12
Figure B10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU (Briaud amp
Gibbens 1994) footings B-13
Figure B11 Characteristic stress versus square root normalized displacements for TAMU footingsB-15
Figure B12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sandB-16
Figure B13 Summary of sand formation factors with corresponding CPT resistancesB-19
Figure B14 Normalized foundation stresses vs square root of normalized displacement for spread
footings on sand (modified after Mayne et al 2012) B-20
Figure B15 Definitions of capacity from three types of observed foundation load test responses
(modified after Kulhawy 2004) B-21
Figure B16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footingsB-22
B-1
Figure B17 Measured load vs displacement curves for 3 shallow foundations on clay at Baytown Texas
(Stuedlein amp Holtz 2010)B-25
Figure B18 Characteristic stress vs square root of normalized displacement for all three foundations at
Baytown site (Stuedlein amp Holtz 2010) B-26
Figure B19 Normalized foundation stress vs square root of normalized displacements B-27
Figure B20 Applied foundation stress vs CPT-calculated stress for 67 footingsB-28
Figure B21 Applied footing stress vs CPT calculated stress on logarithmic scales B-29
Figure B22 Summary graph and equations of direct CPT method for evaluating the vertical B-30
Figure B23 Foundation soil formation parameter hs versus CPT material index Ic B-32
Figure B24 Bearing capacity ratio qmaxqnet versus CPT material index IcB-33
Figure B25 Influence factor for rectangular foundations from elastic theory solution (Mayne amp
Dasenbrock 2017) B-34
List of Tables
Table B1 Direct CPT methods for bearing capacity of footings on clean sandsB-4
Table B2 Direct CPT methods for bearing capacity of footings on clays B-9
Table B3 Summary of large footings soil conditions and reference sources of database B-17
Table B4 Sources for shallow foundation performace B-34
B-2
B1 DIRECT CPT METHOD FOR SHALLOW FOUNDATIONS
B11 Introduction
The analysis of shallow foundations on soils is classically handled as a two-step and two-part set of
calculations involving (a) bearing capacity commonly reliant on limit plasticity solutions and (b)
settlement or more appropriately termed displacements that are assessed via elastic continuum theory
The utilization of cone penetration tests (CPT) provides the necessary geotechnical data for the site-
specific information on the subsurface conditions at the project site including the geostratigraphy and
evaluation of input parameters This is still a viable approach where the CPT readings are interpreted to
give the soil unit weight (γt) effective friction angle (ϕ) undrained shear strength (su) preconsolidation
stress (σp) and elastic moduli (D and E) for analysis
An alternative approach is the use of CPT results to provide direct assessments of bearing capacity
andor settlements A comparison of these two distinctly different and alternate paths is depicted in
Figure B1
Figure B1 Conventional methods versus direct CPT approach to shallow foundation response
A number of available direct CPT approaches for shallow footings are reviewed within this report
Moreover as the entire load-displacement-capacity response of shallow foundations to loading occurs
as a continuous nonlinear phenomenon a single direct CPT solution is presented for the behavior of
B-3
footings on sands (drained) and clays (undrained) The methods discussed herein are specifically to
address the general case of vertical loading of shallow foundations Additional more complex situations
that consider load eccentricity moments inclined forces sloping ground and other facets may be
handled using well-established procedures that are discussed elsewhere (eg Vesić 1975 Kulhawy et
al 1983)
This report focuses on foundations on granular soils andor soils exhibiting drained behavior since
Federal Highway Administration (FHWA) studies have concluded that less than 1 of shallow
foundations for highway bridges are placed on clay soils (Paikowsky et al 2010) One reason for this is
because soft-firm intact clays require a more complicated process that assesses both a short-term
analysis (undrained) as well as long-term analysis (drained) thus requiring undrained bearing capacity
undrained distortion displacements drained bearing capacity and drained primary consolidation
settlements
B12 Direct CPT Methods for Bearing Capacity of Footings on Sands
Classical bearing capacity solutions are based most often in limit plasticity theory albeit can be
formulated from limit equilibrium cavity expansion and numerical methods Because the limit
plasticity solutions are prevalent and adopt total stress analysis they are usually developed as either
fully drained (Δu = 0) or fully undrained (ΔVV = 0) where Δu equals excess porewater pressure and
ΔVV equals volumetric strain Paikowski et al (2010) provides an extensive review of various solutions
There are a number of direct CPT methods for the evaluation of the bearing capacity of footings and
shallow foundations Table B1 lists a variety of these direct CPT approaches for footings on sands In
the direct approach a one-step process is used to scale the measured cone penetrometer readings (ie
measured cone tip resistance (qc) total cone tip resistance (qt) sleeve friction (fs) andor measured
porewater pressure u2)) to obtain the ultimate bearing stress (qult) of the foundation
Table B1 Direct CPT methods for bearing capacity of footings on clean sands
Method Surface Footing Remarks and Embedment Notes
Meyerhof (1956) qult = qc (B12) middotcw qult = qc (B12)(1+DeB)cw
Note for silty sand reduce by 05
cw = 10 dry or moist sand cw = 05 submerged sand
Meyerhof (1974) qult = qc (B + D)40 with stresses in tsf
Presumably dry sands B = fdn width (feet) and D = depth (feet)
Schmertmann (1978)
NA (applies to embedded footings)
Square qult =055σatm(qcσatm)078
Strip qult =036σatm(qcσatm)078
See Figure A2
Embedment applies De gt 05(1+B) for Blt1m De gt 12 m for B gt1m
B-4
Canadian qult = Rk0middotqc Applied to FS = 3 Geotech Society where Rk0 = 03 where FS = factor of (CFEM 199 2) safety
Tand et al (1995) NA qult = Rkmiddotqc + σvo See Figure A3 where Rk = fctn(De B)
Frank and Magnan (1995)
qult = Rk0middotqc where Rk0 = fctn(sand
consistency) Loose Rk0 = 014
Medium Rk0 = 011 Dense Rk0 = 008
qult = Rk1middotqc + σvo where Rk1 = function (De B L amp
sand consistency) Factor Rk1 = Rk0[1+Rk206+04(BL) (DeB)]
and Rk2 = 035 (loose) 050 (medium) and 085 (dense)
Loose sand qc lt 5 MPa
Medium sand 8 MPa lt qc lt 15MPa
Dense sand qc gt 20 MPa
Eslaamizaad amp Robertson (1996)
qult = KΦmiddotqc See Figure A4
See surface footing equation KΦ = function (BDe shape and density)
Lee amp Salgado (2005)
qbL = βbcmiddotqc(AVG)
where qc averaged over distance B
Not addressed See Figure A5 for factor βbc = fctn(B DR
K0 and sB)
beneath the base
Eslami amp Gholami (2005
2006)
See embedded footing solution
qult = Rk1middotqc where Rk1 = function (ratio DeB
and normalized qcσvo) See Figure A6
Measured qc and qcσvo are geometric means over 2B deep
beneath footing
Robertson amp Cabal (2007)
qult = KΦmiddotqc with KΦ = 016
See surface footing equation
Briaud (2007) qult = KΦmiddotqc with KΦ = 023
Based on full-scale tests at Texas AampM
Mayne amp Illingworth
(2010) qult = 018 qc-Mean
Note qc -Mean is averaged CPT cone resistance over depth of
influence z = 15 B deep
Based on 30 footing load tests on 12
sands
Lehane (2013) qult = 016 qc-AVE Note qc-AVE is averaged CPT cone resistance within z (m) = [B (m)
]07
Based on 47 load tests
Table B1 Continued
Notes B = minimum footing width (or diameter) L = footing length De = embedment depth cw = water
table correction σvo = total overburden stress at bearing elevation and σvo = the effective vertical stress
B-5
Figure B2 Direct bearing capacity relationship for embedded square and strip footings
situated on clean sands (after Schmertmann 1978)
Figure B3 Direct CPT bearing factor Rk as function of footing width B and embedment depth
from finite element analyses (Tand et al 1995)
B-6
Figure B4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio
and sand consistency (Eslaamizaad amp Robertson 1996)
B-7
Figure B5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp
Salgado (2005) in terms of footing size relative density and base settlement
Figure B6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and
normalized cone tip resistance (after Eslami amp Gholami 2005)
B13 Direct CPT Methods for Bearing Capacity of Footings on Clays
For direct CPT evaluations of foundation bearing capacity in clays Table B2 summarizes some of the
well-documented methods that are available Generally these assume that the loading is fast enough
and the permeability of the clay is also sufficiently low such that an undrained (ie constant volume)
condition is maintained This is fine for short term loading of foundations particularly for soft to firm
intact clays However the undrained case is not permanent Given sufficient time excess porewater
pressures in the clay will eventually dissipate and hydrostatic conditions will return to equilibrium
During this dissipation phase primary consolidation will occur that results in drained settlements Also
a drained bearing capacity condition will prevail As the CPT is advanced at a constant rate of 20 mms
the recorded readings normally constitute undrained behavior from a direct measurement viewpoint
Thus direct CPT methods are normally focused at an undrained foundation response Nevertheless the
drained footing case can be addressed using CPT results via use of conventional analysis approach and
effective stress limit plasticity solutions that require piezocone penetrometers with porewater pressure
measurements
B-8
B
D)
L
B4060(3501320kc
Many of the earlier solutions were based on data from mechanical penetrometers andor older electric
cones that measure the uncorrected tip resistance (qc) It is now well-recognized that this measurement
must be updated to the corrected cone resistance (qt) because of porewater pressure effects that act on
the mechanics of the load cell and geometry of the particular penetrometer design The results are
particularly significant in clays silts and mixed soils that are soft to firm to stiff where excess porewater
pressures are recorded
Table B2 Direct CPT methods for bearing capacity of footings on clays
Method Footing Comments NotesRemarks
Meyerhof (1974) qult = αbcmiddotqc
where 025 le αbc le 050
Applicable to saturated insensitive clays under short-term loading
Cone tip resistance
measured by
mechanical CPT
Trofimenkov (1974)
Approximation
qult asymp (qc33)09
(Assuming FS = 3)
Strip footings on clays and sandy clays 06m le B le 15m 1m le zemb le 25m
Mechanical CPT
with qc and qult in
kgcm2
Schmertmann (1978)
qult = function (qc and foundation shape)
Relationship shown in Figure A7 for square and strip footings
Cone tip resistance
measured by
mechanical CPT
Tand et al (1986)
qult = Rkmiddot(qc - σvo) + σvo
Factor Rk obtained from Figure A8 accounts for surface and deep footings Separate Rk factor for intact and jointed clays
= (qc1middotqc2)05 qc
where qc1 is geometric mean from bearing elevation to 05B deeper and qc2 is geometric mean from 05B to 15B beneath foundation base
Mix of data from
mechanical CPT qc
and electric CPT qc
LCPC Method
(Frank amp Magnan 1995)
Footing a surface
qult = kc middotqc + σvo
where kc = 032
Embedment case
Note Methods above generally consider undrained bearing capacity
B-9
Figure B7 Direct relationship between ultimate bearing stress in clays and measured cone tip
resistance (Schmertmann 1978)
B-10
Figure B8 Direct relationship between ultimate foundation bearing stress in clays and net
cone tip resistance (after Tand et al 1986)
B14 Direct CPT Assessments of Settlement and Displacements of Shallow Foundations
Methods for evaluating the magnitude of foundation displacements (s) can be found using
elastic theory subgrade reaction methods spring models and numerical simulations The
classical approach to footing settlement calculations on sands is to utilize elastic theory
solutions (eg Poulos amp Davis 1974) which take on the form
119902∙119861∙119868∙(1minus1205842) 119904 = B1
119864119904
where q = applied footing stress and I = displacement influence factor from elasticity theory
The value of I depends upon foundation geometry footing rigidity embedment depth variation of soil
modulus with depth compressible layer thickness and other factors as discussed by Mayne amp Poulos
(1999) For instance for a flexible circular footing of diameter B resting on an infinitely deep soil
formation having a constant modulus Es with depth the influence factor I is equal to 1 for the center
point The results of in-situ field tests such as pressuremeter tests (PMT) standard penetration tests
(SPT) cone penetration tests (CPT) flat dilatometer tests (DMT) and other methods can be used to
ascertain the input value of soil modulus Es
Alternatively direct CPT methods have been investigated to provide a one-step assessment of
foundation displacements Selected methods for direct CPT evaluation of shallow foundation
settlements on granular soils is include DeBeers amp Martens (1957) Meyerhof (1965) DeBeer (1965)
Thomas (1968) Schmertmann (1970) Meyerhof (1974) Berardi amp Jamiolkowski amp Lancellotta (1991)
Robertson (1991) Lutenegger amp DeGroot (1995) Lehane amp Dougherty amp Schneider (2008) Gavin amp
Adekunte amp OrsquoKelly (2009) Mayne amp Illingworth (2010) and Uzielli amp Mayne (2012)
B141 Large Scale Footing Load Tests
The measured load-displacement response of shallow foundations is conducted in the field using either
stepped loading procedures or continuous rate of displacement methods Results from the well-
documented test program at the Texas AampM University (TAMU) site (Briaud amp Gibbens 1999 Briaud
2007) for five spread footings on sand are shown in Figure B9 Each footing clearly shows a nonlinear
B-11
behavior to loading over the range of testing This site is one of the national geotechnical
experimentation sites (NGES) funded by the National Science Foundation (NSF) FHWA and American
Society of Civil Engineering (ASCE)
Figure B9 Measured load-displacements for each of the five large footings at TAMU (Briaud
amp Gibbens 1994) sand site
B142 Generalized Direct CPT Method for Footing Response on Soils
The use of applied foundation stress versus normalized displacement curves (q vs sB) is generalized to
footings on sands silts intact clays and fissured clays A square root plotting of the normalized
displacements (sB) permits a single parameter characterization for each specific soil type that in turn
correlates with the net cone tip resistance The generalization is based on a statistical review of data
from large foundations (B gt 05m) involving 70 full-scale load tests with available CPT data For fine-
grained soils the data are from electronic piezocone tests where the proper correction from qc to qt has
been obtained to allow the best possible results The methodology permits an evaluation of the load-
displacement-capacity response of shallow square footings based on CPTs performed in the four types
B-12
of ground conditions For the TAMU footings the characteristic stress-normalized displacement
response is shown in Figure B10 where all five foundations can be represented by a single curve
Figure B10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU
(Briaud amp Gibbens 1994) footings
B143 Characteristic Stress-Displacement Curves
Fellenius amp Altaee (1994) recommended the concept of a characteristic stress (q) vs normalized
displacement (sB) curve for foundation response on a given soil formation later supported by Decourt
(1999) Briaud amp Gibbens (1999) Lutenegger amp Adams (2003) and Briaud (2007) In this approach the
measured load (Q) vs displacement from individual footings of various sizes that rest on the same soil
conditions collapse to a unified relationship given in Equation B2
119939119943 119956 119954119938119953119953119949119946119942119941 = 119938119943 (119913
) B2
B-13
where af and bf are empirical fitting coefficients (Decourt 1999 Uzielli amp Mayne 2011 2012)
In a review of measured load tests data from large spread footings situated on different sands it has
been suggested that Equation B2 can be reduced to a single parameter expression by use of square root
plotting where the exponent bf is equal to 05 (Mayne amp Illingworth 2010 Mayne et al 2012)
119956 119954119938119953119953119949119946119942119941 = 119955119956radic
119913 B3
where rs = fitted soil parameter from best fit line regression In the case of sands the coefficient rs
represents the supporting soil conditions including particle size relative density and fines content as
well as geologic origin aging and overconsolidation effects
The results from the five TAMU footings are presented in this format in Figure B11 that shows the
formation factor rs = 486 MPa for this sand deposit This single coefficient can be used to express the
nonlinear load versus settlement of any size footing on this sand site Moreover one criterion from
Eurocode for bearing capacity that is used by the European community identifies that stress (or load)
which corresponds to (sB) = 10 That is when the displacement equals ten percent of the footing
width then the capacity has been reached Therefore adopting this criterion for footings on sands
the ultimate bearing stress (qult) for footings on sand can be taken when (sB) equals 010 or when
square root of (sB) equals 0316 In that case this gives qult equal to 0316 middot rs For the TAMU site this
gives qult equal to 0316 times (486) which equals 153 MPa as illustrated by Figure B12
B-14
Figure B11 Characteristic stress versus square root normalized displacements for TAMU
footings
B-15
Figure B12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sand
B15 Footing Database
A database of spread footings and large plates was assembled which considers only full-size shallow
foundations (05 le B le 6 m) that rest on sands silts and clays Sites for the foundations were also subjected to CPT The inclusion solely of large footings are significant because results from small-scale
model footings exhibit scale effects In fact experimental centrifuge work and numerical simulation
studies have shown that bearing capacity factors are size-dependent especially for factor Nγ decreasing
with footing width B (Kimura et al 1985 Cerato amp Lutenegger 2007 Mase amp Hashiguchi 2009) A
direct approach based on full-size footings provides a more reliable means for foundation evaluation
The compiled database is given in Table B3 with a total of 70 large footings and plates These include a
listing of 34 foundations on 13 sands 11 footings on 4 silt deposits 13 footings or large plate load tests
on 6 intact clays and 12 foundations on 5 fissured clays Most of the footings were square or nearly
square (80) while the remaining were circular The largest foundation was a mat 5 m by 14 m in plan
loaded by stacking Kentledge concrete blocks to cause a bearing failure in the underlying soft clay For
sands the largest footing consisted of the 6-m square concrete pad In terms of equivalent square
footings mean footing sizes were 155 m (sands) 114 m (silts) and 126 m (clays)
B-16
Additional details on the individual load tests and site conditions are given elsewhere (Mayne 2009
Mayne amp Illingworth 2010 Uzielli amp Mayne 2011 2012 Mayne et al 2012 Mayne amp Woeller 2014)
Table B3 Summary of large footings soil conditions and reference sources of database
Sand Site Location Soil Conditions Footings Numbers
Shapes and Sizes
ReferencesSource
College Station
Texas Pleistocene sand 5 Square 1 15 25 33 m Briaud amp Gibbens (1999)
Kolbyttemon Sweden Glaciofluvial sand 4 Rect B = 06 12 17 24 m
Bergdahl et al (1985 1986)
Fittija Sweden Glaciofluvial sand 3 Rect B = 06 17 24 m Bergdahl et al (1984 1985)
Alvin West Texas Alluvial sand 2 Circular D = 235 m Tand et al (1994)
Alvin East Texas Alluvial sand 2 Circular D = 22 m Tand et al (1994)
Perth Australia Silceous dune sand 4 Square B = 05 and 10 m
Lehane (2008)
Grabo T1C Sweden Compacted sand fill
1 Square B = 046 m Long (1993)
Grabo T2C Sweden Compacted sand fill
1 Square B = 063 m Long (1993)
Grabo T3C Sweden Compacted sand fill
1 Square B = 080 m Long (1993)
Labenne France Dune sand 6 Square B = 07 and 10 m
Amar et al (1998)
Green Cove Florida Brown silty sand 1 Circular D = 182 m Anderson et al (2006)
Durbin South Africa
White fine sand 1 Square B = 609 m Kantley (1965)
Porto Portugal Residual clayey sands
1 Circular D = 12 m and 1 plate
Viana da Fonseca (2003)
Jossigny France Soft clayey silt 2 Square B = 1 m Amar et al (1998)
Tornhill Sweden Glacial Baltic till 3 Square B = 05 1 2 m Larsson (2001)
Vagverkel Sweden Stiff medium silt 3 Square B = 05 1 2 m Larsson (1997)
Vattahammar Sweden Brn-Gry layered silt 3 Square B = 05 1 2 m Larsson (1997)
B-17
Table B3 Continued
Clay Site Location Soil Conditions Footings Numbers Shapes and Sizes
ReferencesSource
Baytown Texas Fissured Beaumont 2 plates with d = 076 m Stuedlein amp Holtz (2008)
Belfast Ireland Soft clay silt ldquosleechrdquo
1 Square Pad B = 20 m Lehane (Geot Engr 2003)
Bothkennar Scotland Soft silty clay 2 Square Pads B = 22 24 m
Jardine et al (1995)
Bangkok Thailand Soft to stiff clay 4 Square 067 075 090 10 m
Brand et al (1972)
Haga Norway Stiff OC clay 2 Square Footings B = 10 m
Andersen amp Stenhammar (1982)
Rio Grande Brazil Sandy residual clay 3 Square 04 07 and 10 m
Consoli et al (1998)
Shellhaven England Soft estuarine clay Rectangular 5 m by 14 m Schnaid et al (1993)
Texas City A Texas Coast Fissured Beaumont 3 circular plates d = 058 m
Tand et al (1986)
Texas City B1 Texas Coast Fissured Beaumont 2 circular plates d = 058 m
Tand et al (1986)
Texas City B2 Texas Coast Fissured Beaumont 1 circular plate d = 058 m
Tand et al (1986)
Alvin Texas Texas Coast Fissured Beaumont 3 circular plates d = 058 m
Tand et al (1986)
For the series of footings on sands Figure B13 shows the summary of measure formation factors (rs) for
the 34 foundation load tests Also indicated on the graph are the corresponding mean qc values of the
sands averaged over 15middotB beneath the bearing elevations
Note also in sands that the qt and net resistance (qnet) are all very close to the measured qc since (a)
porewater pressure corrections for the a-net value are negligible in clean sands and (b) overburden
stresses are typically only 1 to 3 of the magnitude of qc This is especially true of shallow depths
Therefore for clean sands qnet is approximately qt which is approximately qc
B-18
Figure B13 Summary of sand formation factors with corresponding CPT resistances
As suggested by Briaud (2007) various in-situ measurements can be used to normalize the results from
cone penetration tests in this case Herein the qc from the CPT soundings in the sands are used to
normalize the footing stress axis as presented in Figure B14 It is evident that the results of the load-
displacement response of all 34 footings on 13 different sands can be captured by the characteristic
stress versus normalized displacement with the inclusion of the site-specific cone tip resistance In this
case the mean trend for all footings and sand sites indicates a simple expression shown in Equation B4
119954119943119952119952119957119946119951119944 119956 Clean quartz-silica sands = 120782 120787120790120787radic B4
119954119940 119913
B-19
Figure B14 Normalized foundation stresses vs square root of normalized displacement for
spread footings on sand (modified after Mayne et al 2012)
The results of measured force-displacement curves (Q vs s) from footing load tests are nonlinear and
thus the definition of capacity must be addressed Kulhawy (2004) identified three main curve types
that occur during load testing depicted in Figure B15 The most common response (Type A) shows no
clear peak and is observed in sands and silts as well as slow loading of insensitive clays and fissured
clays In this case ldquocapacityrdquo can be defined by the Eurocode corresponding to the load (or stress) when
sB=10 (Amar et al 1998) For foundations situated on many clays a plateau capacity is reached as
depicted as Type B response In rare cases where sensitive clays or structured soils are encountered a
Type C relationship occurs whereby the load-displacement curve reaches a peak followed by strain
softening In the one case observed in this study for Type C loading (Haga clay) the corresponding
ldquopeak capacityrdquo was reached at a pseudo-strain of sB = 4 as shown in Figure B16 followed by some
strain softening In the cases of soft clays at Belfast and Bothkennar a plunging type (plateau failure)
was achieved at around sB=7
B-20
Figure B15 Definitions of capacity from three types of observed foundation load test
responses (modified after Kulhawy 2004)
B-21
Figure B16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footings
In the case of fine-grained soils that develop excess porewater pressure during cone penetration the
use of the net total cone resistance (qtnet) which equals the cone tip resistance minus the total stress
must be considered (Lunne et al 1997) As noted earlier the correction of CPT data in sands is minimal
because of the low value of penetration porewater pressures and the fact that total overburden stress is
small relative to the cone resistance especially for shallow soundings Thus the use of qtnet may be
substituted for qc into the trend of Figure B1
B16 Undrained and Drained Capacity
The footings on sands typically show drained response with no excess porewater pressures developed
during loading For the foundations situated on silts analyses by Larsson (1997 2001) concluded that
these too show drained conditions For shallow foundations on clays the entire range of drainage cases
are possible ranging from undrained to partially-drained to fully-drained depending upon the applied
rate of loading relative to the permeability of the geomaterial If sufficient data were available for each
of the clay sites then the recent evaluation of drainage condition could be assessed using normalized
velocity criteria (Chung et al 2006) However this was not possible with the current data set Instead
B-22
the test loadings of footings on clays have essentially been assumed to primarily occur under undrained
conditions following recommendations by Tand et al (1986) For this case a limiting bearing stress can
be calculated from bearing capacity analyses and related directly to the CPT readings
From classical bearing capacity calculations using limit plasticity solutions the ultimate foundation stress
is shown in Equation B5
= 119925119940 ∙ 119956119958 B5 119954119958119949119957
where Nc equals the bearing factor for constant volume (514 for strip and 614 for square or circular
plan) and su equals the undrained shear strength of the clay For the case of soft to firm clays of low to
medium sensitivity the operational strength can be related to the preconsolidation stress (Mesri 1975
Jamiolkowski et al 1985) as shown in Equation B6
prime 119956119958 = 120782 120784120784120648119953 B6
Finally the effective preconsolidation stress has been linked directly to the net cone resistance (Chen amp
Mayne 1996 Demers amp Leroueil 2002) as shown in Equation B7
prime 120648119953 = 120782 120785120785(119954119957 minus 120648119959119952) B7
Combining Equations (B5) (B6) and (B7) results in direct expressions for undrained bearing capacity on
intact clays from CPT results as shown in Equations B8 and B8-2
Strip footing 119954119958119949119957 = 120782 120785120789120785(119954119957 minus 120648119959119952) B8
B-23
Squarecircle 119902119906119897119905 = 0445(119902119905 minus 120590119907119900) B8-2
Equation (B8-2) is in excellent agreement with Tand et al (1986) database study for the case of intact
clays and footings with no embedment Their study also provided a lower relationship for foundations
situated on jointed clays specifically recommending that the bearing stress not exceed 030 qtnet
Review of the available data in Figure 3 shows this to be rather conservative and a more realistic value
may be taken at 040 qtnet for fissured clays
B161 Normalized Undrained Footing Response
The characteristic stress vs normalized displacement curves for footings on clays exhibiting undrained
conditions can be verified using a relatively recent case study on Beaumont clay by Stuedlein amp Holtz
(2010) The Baytown Texas site was investigated using an exploration program of soil borings
piezocone soundings and laboratory testing The field testing included a large square footing (B = 276
m) and two small circular plates (D = 076 m) The measured load-displacement responses for all three
foundations are presented in Figure B17 Of course the large footing carried considerably higher loads
than the two small plates on the same soil deposit Yet using the aforementioned characteristic stress
(q) vs square root of normalized displacement (sB) essentially gave a unique relationship for all 3
foundations as presented in Figure B18 In this case utilizing the form of Equation B2 the
characteristic coefficient rs is 177 MPa for this deltaic clay formation
B-24
Figure B17 Measured load vs displacement curves for 3 shallow foundations on clay at
Baytown Texas (Stuedlein amp Holtz 2010)
B-25
Figure B18 Characteristic stress vs square root of normalized displacement for all three
foundations at Baytown site (Stuedlein amp Holtz 2010)
B17 General Direct CPT Method
In a manner analogous to the normalization of applied foundation stresses shown in Figure B14 for
footings on sands a generalized direct CPT method for shallow foundations on different soils can be
made in Equation B9
120782120787 119956 119954 = 119945119956 ∙ 119954119957119951119942119957 ∙ (119913
) lt 119954119940119938119953119938119940119946119957119962 B9
where qcapacity is equal to the foundation bearing capacity of the ground and hs is equal to the empirical
fitting term that depends on soil type Specifically hs equals 058 for sands 112 for silts 147 for
fissured fine-grained soils and 270 for clays A summary of the normalized stress versus the normalized
displacement curves for these four soil types is presented in Figure B19 where the data fitting to obtain
the hs parameter on clays are limited to (sB) lt 004 corresponding to undrained loading The data on
B-26
silts and sands are considered fully drained whereas the fissured clay subset may be partially drained to
undrained The statistical measures for obtaining the fitted parameter hs are quite good as shown in
Figure B20 with the coefficient of determinations (r2) values of 0947 for sands 088 for silts 0935 for
fissured and 0925 for intact clays Additional statistical evaluations on the database are provided by
Uzielli amp Mayne (2011)
Figure B19 Normalized foundation stress vs square root of normalized displacements
B-27
Figure B20 Applied foundation stress vs CPT-calculated stress for 67 footings
The same data are shown on Figure B21 in a log-log plot to emphasize the early parts of the load-
displacement responses of these footings The best fit line from regression analyses shows excellent
statistical correlations As detailed earlier the undrained bearing capacity of square or circular footings
on clays can be taken as approximately 040 times qtnet For silts and sands that experience drained
loading with no excess porewater pressures being developed and no clear peak value for capacity the
(sB) =10 criterion can be used In this case Equation 7 can be used directly to obtain stress q equal to
qcapacity when (sB)05 is 0316
An alternate approach is presented in Figure B22 summarizing the direct CPT method for estimating the
load-displacement-capacity of shallow square footings on four soil types The maximum values of soil
bearing capacity (qmax) can be capped at stresses corresponding to a percentage of the average
measured qtnet determined over the depth interval from the foundation bearing elevation to 15middotB
below the foundation In such an approach the foundation bearing capacity can be taken as
(qcapacityqtnet) of 02 for sands 035 for silts 040 for fissured and 045 for intact clays Using the assigned
values of the hs parameter the corresponding cut-off for the general stress-normalized displacement
relationship given by Equation 7 can be established specifically giving corresponding values of the
B-28
normalized displacements which can also be considered as pseudo-strains equal to (sB) max of 4 for
clays 7 for fissured clays 10 for silts and 12 for sands
Figure B21 Applied footing stress vs CPT calculated stress on logarithmic scales
B171 CPT Soil Material Index Ic
The soil behavioral type can be assessed indirectly by CPT using a material index (Ic) as defined by
Robertson (2009a b) shown in Equation B10
119920119940 = radic[120785 120786120789 minus 119949119952119944119928119957119951]120784 + [120783 120784120784 + 119949119952119944119917119955]120784 B10
B-29
Figure B22 Summary graph and equations of direct CPT method for evaluating the vertical
where Qtn equals the stress-normalized cone tip resistance and Fr equals the normalized sleeve friction
calculated from the cone penetrometer readings as shown in Equations B11 and B12 respectably
(119954119957minus120648119959119952)120648119938119957119950 = 119951 B11 119928119957119951 prime (120648119959119952120648119938119957119950)
119943119956 119917119955() = 120783120782120782 ∙ B12 (119954119957minus120648119959119952)
where σatm equals a reference stress equal to atmospheric pressure (σatm = 1 atm asymp 1 bar asymp 100 kPa) In
the initial evaluation the exponent n is set to n = 1 to find the soil behavioral type (SBT) based on a 9-
zonal chart as shown in Figure 23 The value of exponent n varies with soil type ranging from n = 1 in
intact clays and decreasing with increasing grain size to around approximately 075 in silts and
B-30
approximately 05 plusmn 02 in clean sands The appropriate value of exponent n is found by iteration until
conversion using the relationship (Robertson 2009b) shown in Equation B13
prime 120648119959119952 119951 = 120782 120785120790120783 ∙ 119920119940 + 120782 120782120787 ( ) minus 120782 120783120787 le 120783 120782 B13 120648119938119957119950
As the distinction that footings on intact clays subjected to fast loading are behaving primarily under
undrained conditions (ie constant volume) Robertson (2009a) suggested that this occurs when Ic gt 27
while in contrast Ic lt 25 corresponds more or less to drained loading (ie no excess porewater
pressures)
The CPT material index can be used to identify soil type and the corresponding characteristic hs value for
estimating footing load-displacement response In a number of the case studies investigated the results
of the sleeve friction readings were also available to allow the calculation of Ic with depth at these sites
Figure B23 shows the tentative relationship between the values of the hs parameter plotted versus the
corresponding CPT material index with an approximate trend given by Equation B14
120784120785 119945119956 = 120784 120790 minus 120783120787 B14
119920119940 120783+( ) 120784120786
B-31
Figure B23 Foundation soil formation parameter hs versus CPT material index Ic
Soil behavioral type from the Ic ranges given by Robertson (2009b) and are shown in Figure B23 with an
overall general agreement with the known soil classifications In summary the CPT can provide an
evaluation of the load-displacement-capacity response directly via Equation B15 which may be of
interest in mixed soil types such as silty sands sandy silts and the like
120784120785 Footing stress 119954 = 119954119951119942119957 ∙ radic(119956frasl119913) ∙ [120784 120790 minus
)120783120787] B15
120783+(119920119940frasl120784120786
As discussed earlier foundation bearing capacity may be taken by a limiting value of pseudo-strain (sB)
max or by qmax as specified as a percentage of qtnet This definition of bearing capacity can be seen in
comparison to soil type as seen in Figure B24
B-32
Figure B24 Bearing capacity ratio qmaxqnet versus CPT material index Ic
B172 Rectangular Foundations
This same elastic solution from Giroud (1968) for a CPT method for square and circular foundations were
utilized for rectangular foundations The study covered rectangular distortions (AB) ranging from 1
(square) to very long foundations with AB equal to 20 where A represents the foundation length and B
represents the foundation width (Mayne amp Dasenbrock 2017) The influence factor for rectangular
shaped foundations is given in Figure B25 and the expression in Equation B16
= (119912119913)120782120785120786120787 119920119912119913 B16
B-33
Figure B25 Influence factor for rectangular foundations from elastic theory solution (Mayne
amp Dasenbrock 2017)
Including 32 full scale load tests on sands results from an additional 98 shallow foundations have been
reported in previous studies The foundations were primarily square or circular and among these include
large spread footings with measured settlements and bearing capacity Table B4 shows a summary of
the number of footings and footing sizes used in the study The range of length to width ratios varied
from 1 lt (AB) lt 23 with an average value of (AB) equal to 238 The embedment depth to footing
width ranged from 0 lt DeB lt 222 and averaged 046
Table B4 Sources for shallow foundation performace
Source of Data No of
Footings
Mean B
(m)
Max B
(m)
Min B
(m)
Mean A
(m)
Max A
(m)
Min A
(m)
Mayne et al (2012) 32 149 609 046 149 609 046
Lehane (2011) 8 041 027 060 041 060 027
Gifford et al (1987) 17 846 1588 527 1591 3596 701
Jeyapalan amp Boehm
(1986)
13 1022 2892 400 1671 3049 640
B-34
Schmertmann
(1970)
31 945 5608 061 1288 8668 061
Papadopoulos
(1992)
29 870 3600 100 1300 7290 100
Total 130 671 5608 027 1012 8668 027
The solution for shallow foundation settlements given back in Equation B1 combined with
the expression in Equation B16 takes on the form
119954∙119913∙119920119912119913∙(120783minus120642120784) 119956 = B17
119916119956
Combining the direct CPT equation developed from the 32 full scale load tests (Equation B15)
together with the influence factor for rectangular shaped foundations becomes
minus120782120785120786120787 120784120785 119912 Footing stress 119954 = 119954119951119942119957 ∙ radic(119956frasl119913) ∙ [120784 120790 minus
)120783120787] ∙ [ ] B18
120783+(119920119940frasl120784120786 119913
B18 Conclusions
A direct CPT method for square rectangular and circular shallow footings is developed using a database
of 166 full-scale field load tests where the minimum footing size of an equivalent square width of 05 m
is established to avoid scaling issues The use of load vs displacement is generalized by characteristic
stress vs normalized displacement (sB) and further simplified by a square root plotting technique that
captures the ground response in a single parameter that is related to the qtnet Data are grouped
according to four main soil categories sands silts fissured clays and intact clays It is generally believed
that the footings on sands and silts exhibit fully drained behavior while in intact clays an undrained
response occurs under conditions of constant volume
B-35
The summary equation for evaluating the vertical stress-displacement-capacity of square rectangular
and circular footings is given by
minus120782120785120786120787 119956 119912 119954119943119952119952119957119946119951119944 = 119945119956 ∙ 119954119957119951119942119957 ∙ radic ∙ ( ) lt 119954119950119938119961 B19
119913 119913
where the parameter hs also relates directly to Ic The foundation capacity depends upon the mode of
failure as well as drainage conditions and can be taken as a fraction of qtnet where qmaxqtnet is 020 in
sands 035 in silts 040 in fissured clays and 045 in intact clays as shown in Figure 24 Alternatively the
capacity can be defined using a limiting pseudo-strain given by (sB) max of 4 in clays 7 in fissured
10 in silts and 12 in sands
B-36
APPENDIX C
List of Figures
Figure C1 Components of Axial Pile Capacity C-3
Figure C2 Selected alpha relationships as function of the undrained shear strengthC-5
Figure C3 Pile side friction coefficient β in terms of effective friction angle and overconsolidation ratio
C-11
Figure C4 Concept of Direct versus Rational Method for CPT evaluation of axial pile capacityC-14
Figure C5 Soil behavior type using CPT via original UniConeC-19
Figure C6 Soil behavior type using CPT via Modified UniCone C-19
Figure C7 Nine-zone soil behavioral soil type using normalized piezocone parameters C-21
Figure C8 Summary of Modified UniCone Method for direct CPT assessment of axial pile capacity C-22
Figure C9 Elastic continuum solution for axial pile response under compression and tension loadingC-24
List of Tables
Table C1 Alpha Methods for Axial Pile Side Friction where fp = αmiddotsu C-5
Table C2 Beta Methods for Axial Pile Side Friction where fp = βmiddotσvoC-9
Table C3 Selection of Direct CPT Methods for Axial Pile CapacityC-15
C-1
C1 EVALUATING DEEP FOUNDATION RESPONSE FROM CONE
PENETRATION TESTS
C11 Introduction
The axial response of deep foundations includes the load-displacement-capacity and axial load transfer
when driven pilings and drilled shafts are loaded in compression and uplift The use of cone penetration
testing (CPT) especially seismic piezocone tests (SCPTu) are advantageous since they provide
information on the subsurface soils and their geomaterial properties The SCPTu provides at least four
measurements on soil behavior with depth including (a) cone tip resistance qt (b) sleeve friction fs (c)
penetration porewater pressure u2 and (d) shear wave velocity Vs This offers the opportunity to
determine the geostratigraphy unit weight effective overburden stress shear strength stress state
and stiffness of the ground from a single exploratory sounding thus values in economic expediency
and reliability The axial pile capacity can be calculated based on static equilibrium of forces acting along
the sides and base of the pile foundation The displacement of the pile can be ascertained using
elasticity theory from closed-form solutions boundary elements andor finite element analyses Load
transfer occurs along the length of the pile and only a portion of axial forces are transmitted to the base
or toe or tip of the pile These too can be assessed using elasticity solutions
An alternate approach to pile capacity involves the utilization of direct CPT methods available
since the 1970s but in the past 10+ years a number of new and statistically reliable algorithms
have been developed which can be implemented for highway design and construction
C111 Axial Pile Capacity
The axial compression capacity of a single pile foundation is composed of a shaft or side component and
end-bearing component at the base as depicted in Figure 1 For a circular pile the side capacity (Qs) is
determined from the unit side friction (fp) acting along the surface area of the shaft which is As = πmiddotdmiddotL
where d = pile diameter and L = length embedded below grade
If the magnitude of side friction is uniform and constant with depth the side capacity is simply
C-2
119928119956 = 119943119953 middot 119912119956 C1
Moreover however many piles extend through multiple layers and a summation of unit side
frictions acting on various pile segments must be tabulated over the length of the pile as
suggested by Figure C1
Figure C1 Components of Axial Pile Capacity
The unit-end bearing resistance (qb) acts over the base of the pile tip where the area of a circular pile is
given by Ab = πmiddotd24 For piles in compression loading the base capacity is determined from Equation
C2
119876119887 = 119902119887 middot 119860119887 C2
and for piles in tension (or uplift) it is normally taken that Qb = 0
C-3
C112 Pile Unit Side Friction
Several different approaches can be adopted for evaluating the unit pile side friction (fp) prior to full-
scale load testing and construction (Poulos amp Davis 1980 ONeill 2001) The most common types
including the alpha and beta methods as summarized in Tables 1 and 2 respectively In the alpha
method an empirical coefficient (α) is applied to the undrained shear strength (su) of clay soils to
obtain
119891119901 = 120572 middot 119904119906 C3
An illustrative example of alpha curves is shown in Figure C2 A difficulty with the alpha
method is the evolution of many variants and changes to the expressions for its estimation It is
also restricted in its use for specific types of piles in clays and fine-grained soils
C-4
Figure C2 Selected alpha relationships as function of the undrained shear strength
Table C1 Alpha Methods for Axial Pile Side Friction where fp = αmiddotsu
C-5
Reference
Source
Alpha Equation Remarks
Tomlinson
(1957)
2716801102
uu ssAll pile types (steel
concrete timber)
Note su in ksf
Tomlinson
(1957)
2616201102
uu ssConcrete piles
Note su in ksf
McClelland
(1974)
Kerisel α = 07 - 031middotln(su)
Peck α = 10 - 02middotsu
Woodward α = 091middot(su)091
Driven piles in clay
Note su in ksf
Semple
(1980) )1(511
)1(1
OCR
OCR Summary from 9 series of
pile load tests
American
Petroleum
Institute (API
1981)
For suσvo le 035 α = 10
For suσvo gt 080 α = 05
Otherwise α = 1 + 1111middot (035 - suσvo)
Driven steel pipe piles
Tomlinson
(1986)
1 α = 55 (su)-091
2 α = 785 (su)-102
3 α = 605 (su)-100
where 75 lt su (kPa) lt 210
1 Driven piles
2 Bored piles
3 Driven piles in till with L
lt 10middotd
API (1987) 1 α = 10 for su lt 25 kPa Driven piles in clay other
than Gulf of Mexico
C-6
2 α = 125 - 001middotsu for 25 lt su lt 75 kPa
3 α = 050 for su gt 75 kPa
Kulhawy amp
Jackson
(1989)
α = 021+026middot(σatm su) le 1 Analyses of 106 drilled
shaft foundations
Table C1 Continued
API (1989) 05 For suσ vo le 1 α =
su radic frasl prime σvo
minus025 su For suσ vo gt 1 α = 05 ∙ ( frasl prime ) σvo
Driven steel pipe piles
Kulhawy amp
Jackson
(1989)
OCR
OCR
sin50
tan)sin1( sin
Λ = (1-CsCc) asymp 080 for
insensitive clays
Nowacki et al
(1996)
minus05 su For suσvo le 07 α = 05 ∙ ( frasl prime ) σvo
minus02 su For suσvo gt 07 α = 055 ∙ ( frasl prime ) σvo
Driven pile foundations
Kolk and van
der Velde
(1996)
α =09middot FL middot (suσvo)-03 lt 10
FL = [ (L - z)d]-02 = length term
L = pile length
Note su obtained from
UU lab tests
Applicable to driven piles
in clays
C-7
Knappett amp α = 1 for su le 30 kPa Non-displacement piles in
Craig (2012) fine-grained soils α = 116 - (su185) for 30 kPa le su le 150 kPa
α = 035 for su gt 150 kPa
Knappett amp
Craig (2012) α = 055 ∙ 02 40
( ) (LfraslD)
∙ su ( frasl prime σvo
minus03
) Displacement piles in fine-
grained soils
z = depth at point considered
d = outside diameter of pile
Karlsrud et al
(2005 NGI
Method)
α = function (suσvo and plasticity index) Driven pilings
Fleming et al
(2009) 05 minus05 su su For su σvo le 1 120572 = ( frasl prime ) ∙ ( frasl prime )
σvo NC σvo
05 minus025 su su For su σvo gt 1 α = ( frasl prime ) ∙ ( frasl prime ) σvo NC σvo
For driven piles in clay
where (suσvo)NC is the
normalized shear strength
for normally-consolidated
clay
Brown et al
(2010)
α = 0 for 0 lt z le 5 feet
α = 055 for z gt 5 feet and (suσatm) le 15
α = 055-01middot(suσatm -15) for 15le (suσatm) le 25
Drilled Shafts and Bored
Piles
Table C1 Continued
C-8
OCRK )sin1(0
Karlsrud
(2012)
α = function (su σvo and plasticity index) Driven pilings
Note as reported by Karlsrud (2012)
In the beta method the coefficient β is applied to the effective overburden stress (σvo) at the point of
concern along the pile length
prime = 120573 C4 119891119901 middot 120590119907119900
Table C2 Beta Methods for Axial Pile Side Friction where fp = βmiddotσvo
Reference Source Beta Equation Remarks
Burland (1973) β = (1-sinϕ) middot tan ϕ Piles in NC clays
Meyerhof (1976) β = Kmiddottanδ where K = K0 = for NC clays
K = 15middotK0 for OC clays
Poulos amp Davis
(1980)
β = (1-sinϕ) middot tan ϕ middot OCR05 Piles in OC stiff clays
Table C2 Continued
C-9
Kulhawy amp Jackson
(1989)
β = (1-sinϕ) middot tan ϕ middot OCRsinϕ Piles in quartz sands and insensitive
clays
ONeill (2001) β = (1-sinϕ) middot tan θ middot OCRsinϕ Drilled shafts where θ = (δϕ)middotϕ and
δ = interface friction between pile and soil
Fleming et al
(2009)
β = Kmiddottanδ K = lateral stress coefficient and δ =
friction angle between soil and pile
material
Karlsrud (2012) β = fctn (OCR and PI) PI = plasticity index of the clay ()
Mayne and Niazi
(2017)
β = CM middot CK middot K0 middot tanϕ
where K0 = (1-sinϕ) middot OCRsinϕ for soils
that are virgin loaded then unloaded
CM = pile material factor = 1 (drilled
augered) 09 (prestressed or precast
concrete) 08 (timber) and 07
(steel) and CK = pile installation factor
= 09 (bored or augered) 10 (low
displacement eg H-pile or open-end
pipe) and 11 (driven high
displacement eg prestressed
concrete closed-end pipe)
As the beta approach applies to all types of soils (gravels sands silts and clays) and more or
less has remained unchanged since its advent circa 1970 it has been selected for further
discussion herein In the direct form for calculation of unit pile side friction the expression is
given by
119874119862119877119904119894119899120601 prime prime 119891119901 = 119862119872 ∙ 119862119870 ∙ (1 minus 119904119894119899120601prime) middot ∙ 119905119886119899120601prime ∙ 120590119907119900 C5
C-10
where CM is a soil-pile interface friction coefficient and CK = pile installation factor Values of CM are in
the range 07 lt CM lt 10 depending upon pile material (steel wood concrete) and ranges for CK vary
09 lt CK lt 11 depending upon method of installation (auger drill driven) as detailed in Table 2
The value of K0 is limited to the passive stress coefficient which for the simple Rankine case is given by
KP = (1+sin ϕ) (1 - sin ϕ ) Thus there is a maximum value of overconsolidation ratio for which
Equation C5 applies given by OCRlimit = [(1+sin ϕ ) (1 - sin ϕ )2] (1sinϕ)
The corresponding graph in Figure C3 shows the relationship for β in terms of ϕ and OCR
Figure C3 Pile side friction coefficient β in terms of effective friction angle and overconsolidation ratio
C113 Pile Unit End Bearing
C1131 Theoretical Considerations
The evaluation of the end bearing resistance of pile foundations is commonly determined using
limit plasticity theory where
prime Drained loading = C6 119902119887 119873119902 middot 120590119907119900
C-11
Undrained loading 119902119887 = 119873119888 middot 119904119906 C7
where the value of σvo is calculated at the depth z = L and the value of su is the average undrained shear
strength from z = L to a depth z = L + d beneath the pile tip For a circular pile the limit plasticity solution
for undrained loading gives (Vesic 1977) a value Nc = 933 while a deep strip foundation would employ a
value of Nc = 824 For drained loading the expression for Nq for a deep foundation is given by (Vesic
1977)
)]arctan()sin1(tan21[)](tan1[sin1
sin1)tanexp( 2 BLABNq
C8
where A and B are the pile plan dimensions (for a circular pile A = B) and L = pile length For a
circular pile this can be approximated by
asymp 077(120601prime75deg) 119873119902 C9
over a range of effective friction angles 20deg le ϕ le 45deg
C1132 Practical Considerations
For drained loading of sands the full calculated end bearing capacity will never be realized because it
would require the pile to move a distance equal to its diameter This is beyond practical use and
therefore the limit plasticity solutions must be clipped to a fraction of the calculated value Another
reason for using a reduced end-bearing capacity is due to strain incompatibility since the side capacity is
mobilized early while the end-bearing is engaged much later So to have compatible values of Qs and
Qb the qb must be reduced using the following guidelines (Randolph 2003 Mayne 2007)
prime prime C10 119902119887 = 119873119902 ∙ 120590119907119900 ∙ 119891119909
where fx = strain incompatibility factor
fx = 010 for drilled shafts augered cast-in-place and bored piles
fx = 020 for driven low displacement piles (opened-ended steel pipe and H-piles)
fx = 030 for driven high-displacement piles (ie solid piles PSC and closed-ended pipe)
C-12
For undrained loading of circular piles in compression no reduction of end-bearing is necessary
therefore
119902119887 = 933 middot 119904119906 C11
C12 Direct CPT Methods for Pile Capacity
In the direct CPT method the penetrometer readings are scaled directly via specified algorithms to
obtain the pile unit side friction and end-bearing as depicted in Figure C4 As many as 40 different
direct CPT methods have been developed over the past five decades as summarized by Niazi amp Mayne
(2013) Starting circa 1970 many of these early methods relied on hand-recorded data from field
mechanical CPTs where only qc data were obtained at 20 cm intervals or later with mechanical readings
of both qc and fs using special sets of inner and outer rods that recorded vertical load for tip and loads
for tip plus sleeve in alternating increments Also early pile load test data were often obtained from
top-down measurements of load-displacement
C-13
Undrained qb = Nc su
Qside = S (fp DAs)
QTotal = Qs + Qb - Wp
fp = cmck Ko svorsquo tanrsquo
Qbase = qb Ab
Drained qb = Nq svorsquo
Method Two Rational or ldquoIndirectrdquo Method
Method OneldquoDirectrdquo CPT Method
(Scaled Pile)
fp = fctn (soil type pile
type qt fs and u2)
qb = fctn (soil type qt-u2)
qb = unit end bearing
unit sidefriction fp
OCR su Ko t DR rsquo
AXIAL PILE CAPACITY FROM CPT READINGS
Figure C4 Concept of Direct versus Rational Method for CPT evaluation of axial pile capacity
Beginning in the mid-1990s as the modern electric piezocone (CPTu) was implemented the newer
equipment offered improved data and better resolution because of the additional reading of porewater
pressures and correction of raw measured qc to total resistance qt In addition the use of electronic
digital data collection and field computers proved superior in field measurements and recordings As a
consequence several reliable CPT methods for axial pile capacity have been developed Moreover
parallel improvements in full-scale pile load testing occurred and now include modern strain gage
instrumentation digital data recording and automated testing procedures Also the testing can be
single direction (compression or tension) or bi-directional as in the Osterberg cell
C-14
Table C3 provides a selection of recent direct CPT methods that have become available over the
past two decades A number of these (namely ICP NGI UWA and Fugro) were funded by the
offshore industry because of the growth of oil amp gas reserves and windfarm installations thus
necessitating increased concerns on risk probability and reliability in the site investigations for
offshore platforms and design of large driven monopile foundations These direct CPT methods
are often statistically based on large datasets compiled from full-scale load tests made
worldwide (eg Schneider et al 2008)
Table C3 Selection of Direct CPT Methods for Axial Pile Capacity
Method Pile Types Soil Types References CPT
data
Additional
parameters
needed
Unicone Method all types Sands silts
clays
Eslami and
Fellenius
(1997)
Fellenius
(2009)
qt fs u2
KTRI = Kajima
Technical
Research
Institute
all types Sands
mixed clays
Takesue et al
(1998)
fs and u2 No guidance given
on end bearing
resistance
Imperial College
Procedure (ICP)
OE and CE Sands Chow et al
(1997 PhD)
Jardine et al
(2005)
qt Interface friction
angle (δ) from
ring shear tests
Correlated to
mean grain size
(D50)
C-15
OE and CE Clays Jardine et al
(2005)
qt Interface friction
angle (δ)
correlated to PI
Table C3 Continued
NGI Method
(Norwegian
Geotechnical
Institute)
OE and CE Sands Clausen et al
(2005)
qt DR from qt
Clays a Alpha
method
(Karlsrud et al
2005 2012)
b Beta
method
(Karlsrud
2012)
qt for su
qt for
OCR
PI = plasticity
index ()
PI = plasticity
index ()
Fugro Method OE and CE Sands Kolk et al
(2005)
qt
Clays Van Dijk and
Kolk (2011)
qt
OE and CE Sands Lehane et al
(2005)
qt IFR = infilling ratio
related to amount
of plugging
C-16
Purdue LFRD Drilled
shafts
Sands Basu amp
Salgado
(2012)
qt LRFD = load
resistance
factored design
Enhanced Various Sands silts Niazi and qt u2 Based on 330 load
Unicone Method clays and Mayne (2015 and fs tests including
mixed soils 2016) bored augered
jacked and driven
piles
UWA (Univ of
Western
Ra = pile
roughness
Australia) Clays Lehane et al
(2012)
qt Interface friction
angle (δ) from
ring shear tests
and also method
without δ
HKU Method
(Hong Kong
Univ)
OE and CE
(end
resistance
only)
Sands Yu and Yang
(2012)
qt End bearing only
PLR = plug length
ratio
Note PLR can be
estimated from
OE inner diam
Table C3 Continued
Note previously called Marine Technical Directorate (MTD)
C-17
Many of the offshore CPT methods relate to driven piles either in sand or clay as that is
common deep foundation for those purposes Of particular interest are the Unicone and
Modified Unicone Methods since they use all three readings of the piezocone (qt fs and u2)
and address a variety of pile foundation types
C13 Modified UniCone Method
The modified UniCone Method was developed based on a total 330 pile load tests which were
associated with SCPTu data during their site investigations (Niazi and Mayne 2015 2016) This
represents a threefold increase over the original UniCone database that was built upon data
from 106 pile load tests (Eslami amp Fellenius 1997)
For the original UniCone algorithms use is made of the effective cone resistance (qE)
119902119864 = 119902119905 minus 1199062 C12
and a chart of qE vs fs provided an approximate soil classification in five distinct groups as shown by
Figure C5 Later in the modified approach a better delineation of the larger dataset gave soil
subclassifications as indicated by Figure C6
C-18
Figure C5 Soil behavior type using CPT via original UniCone
Figure C6 Soil behavior type using CPT via Modified UniCone
C-19
In the modified approach the 9-zone normalized soil behavioral type (SBTn) is ascertained using CPT
data in conjunction with the Robertson (2009) charts as shown in Figure C7 As discussed in Appendix
A and B the SBTn method uses the normalized cone resistance (Qtn) normalized sleeve friction (Fr())
and CPT material index Ic This permits a much wider range in the calculated pile side friction because fp
is related as a continuous curve with Ic rather than only 5 values that are assigned in the original
scheme
The pile unit side friction (fp) is obtained from qE and the CPT material index Ic using the following
expression at each elevation along the sides of the pile
10(0732 middot 119868119888 minus 3605) fp middot middot C13 = 119902119864 middot 120579119875119879 middot 120579119879119862 120579119877119860119879119864
C-20
Figure C7 Nine-zone soil behavioral soil type using normalized piezocone parameters
(after Robertson 2009 Mayne 2014)
where θPT = coefficient for pile type (θPT = 084 for bored piles 102 for jacked piles 113 for driven
piles) θTC = coefficient for loading direction (θTC = 111 for compression and 085 for tension) and θRATE
= rate coefficient applied to soils in SBT zones 1 through 7 (θRATE = 109 for constant rate of penetration
tests and 097 for maintained load tests) Since CPT provides data at regular intervals of 2 cm to 5 cm
along the sides of the pile the average fp from z = 0 to z = L can be used directly in Equation 1 to obtain
the shaft capacity Qs
C-21
The pile end bearing resistance is obtained from
middot 10(0325middot 119868119888minus1218) 119902119887 = 119902119864 C14
where qE is averaged in the vicinity of the pile tip Figure C8 shows the Modified Unicone Method in
graphical format For sensitive clays of zone 1 please see additional discussions by Niazi amp Mayne
(2016)
Figure C8 Summary of Modified UniCone Method for direct CPT assessment of axial pile
capacity
C-22
C14 Axial Pile Displacement
The movement of pile foundations can be assessed using elastic continuum theory (Poulos amp
Davis 1980 Randolph 2003 Mayne amp Niazi 2017) These relationships have been developed
using finite element analyses boundary elements and analytical closed-form solutions For the
latter approach the top displacement of a rigid pile subjected to a vertical force is shown in
Figure C9 This gives the movement at the top of a pile subjected to either compression or
tension (uplift) loading Also the percentage of axial load transferred to the pile toe is
determined For piles extending through various soil layers the elastic solution can be
implemented by using a set of stacked pile segments each with its own stiffness as
represented by a soil Youngs modulus
For pile groups the use of computer software is recommended Several available programs can handle
pile groups under axial and lateral moment loading such as DEFPIG (Univ Sydney) and PIGLET (Univ
Western Australia) A full listing of pile foundation software is given at the Geotechnical amp
GeoEnvironmental Service Directory wwwggsdcom
Ps
= Sh
aft
Load
sL
t
tEd
IPw
Load Transfer to Base
)]1)((5ln[
)(
)1(1
1
2 vdL
dLFI
E
ECT
211
IF
P
P CT
t
b
Base Load = Pb
Randolph Elastic Solution
Ground Surface
Top Load Pt = Ps + Pb
Rigid Pile Response
wt = displacementd = diameter
L = length
Es = Elastic Soil Modulus
Es0 (at z = 0) surface
EsM (at z = L2)mid-length
EsL (at z = L)full length
Load Transfer to Shaft
E = EsMEsL = Gibson parameterE = 1 for homogeneous case (constant E)E = 05 for pure Gibson case (Es0 = 0)
where FCT = load direction (= 1 compression 0 tension)
21
IF
P
P CT
t
b
z = depth
= Poissons ratio
Tension
Compression
C-23
Figure C9 Elastic continuum solution for axial pile response under compression and tension
loading
C-24
C15 Acknowledgements
The author appreciates the help of Fernando Illingworth of PonsEcuador Meena Viswanth of
GeoSyntecAtlanta and Dr Marco Uzielli of GeoRiskItaly in helping to collect and process the data
Funding support was provided by David Woeller of ConeTec Richmond BC
C-25
- Structure Bookmarks
-
- CONE PENETRATION TEST DESIGN GUIDE FOR STATE GEOTECHNICAL ENGINEERS
- TABLE OF CONTENTS
- LIST OF TABLES
- CHAPTER 1 INTRODUCTION
- CHAPTER 2 DIRECT CPT METHOD FOR SOIL CHARACTERIZATION
- CHAPTER 3 DIRECT CPT METHOD FOR SHALLOW FOUNDATIONS
- CHAPTER 4 DIRECT CPT METHOD FOR DEEP FOUNDATIONS
- REFERENCES
- APPENDIX A
- APPENDIX B
- APPENDIX C
-
CONE PENETRATION TEST DESIGN GUIDE FOR STATE
GEOTECHNICAL ENGINEERS
Prepared by
Ryan Dagger
David Saftner
Department of Civil Engineering
University of Minnesota Duluth
Paul W Mayne
School of Civil amp Environmental Engineering
Georgia Institute of Technology Atlanta
November 2018
Published by
Minnesota Department of Transportation
Research Services amp Library
395 John Ireland Boulevard MS 330
St Paul Minnesota 55155-1899
This report represents the results of research conducted by the authors and does not necessarily represent the views or policies
of the Minnesota Department of Transportation the University of Minnesota Duluth or the Georgia Institute of Technology
This report does not contain a standard or specified technique
The authors the Minnesota Department of Transportation the University of Minnesota Duluth and the Georgia Institute of
Technology do not endorse products or manufacturers Trade or manufacturersrsquo names appear herein solely because they are
considered essential to this report because they are considered essential to this report
TABLE OF CONTENTS
CHAPTER 1 INTRODUCTION1
CHAPTER 2 DIRECT CPT METHOD FOR SOIL CHARACTERIZATION 4
21 Introduction 4
22 Soil Unit Weight 6
23 CPT Material Index 7
231 Step 1 Normalized Sleeve Friction 7
232 Step 2 Iteration 8
24 Soil Behavior Type (SBT) 8
241 Step 1 8
242 Step 2 9
25 Effective Stress Friction Angle 9
26 Stress History 10
27 Lateral Stress Coefficient 11
28 Undrained Shear Strength 12
29 Ground Stiffness and Soil Moduli 12
210 Coefficient of Consolidation 13
211 Hydraulic Conductivity14
212 Example Problems 15
2121 Example 1 Direct CPT Methods for Geoparameters on Sands 15
2122 Example 2 Direct CPT Methods for Geoparameters on Clay 33
CHAPTER 3 DIRECT CPT METHOD FOR SHALLOW FOUNDATIONS 56
31 Procedure 56
311 Step 1 Estimating Footing Dimensions57
312 Step 2 Soil Characterization 58
313 Step 2a Foundation soil formation parameter59
314 Step 3 Soil elastic modulus and Poissonrsquos ratio 60
315 Step 4 Net cone tip resistance 60
316 Step 5 Bearing capacity of the soil 61
317 Step 6 Settlement61
318 Step 7 Final Check 61
32 Example Problems 62
321 Example 3 Direct CPT Method on Sands 62
CHAPTER 4 DIRECT CPT METHOD FOR DEEP FOUNDATIONS 70
41 Introduction70
42 Modified UniCone Method72
421 Step 172
422 Step 272
423 Step 373
424 Step 473
43 Axial Pile Displacements 73
44 Example Problems 74
441 Example 5 Direct CPT Methods Axial Pile Capacity74
REFERENCES 87
APPENDIX A
APPENDIX B
APPENDIX C
1
2
3
4
5
6
7
8
9
10
11
12
13
14
15
16
17
18
19
20
21
22
23
24
LIST OF FIGURES
Figure Geoparameters determined from CPT 1
Figure Conventional method for shallow foundation design compared to direct CPT method 2
Figure Direct CPT evaluation of axial pile capacity 3
Figure Soil unit weight from CPT sleeve friction 6
Figure Approximate ldquorules of thumbrdquo method using CPT sounding from Wakota Bridge MN 7
Figure SBT zones using CPT Ic 9
Figure Yield stress exponent compared to CPT material index 11
Figure k vs dissipation time for 50 consolidation (Mayne 2017) 15
Figure CPT data from Benton County Minnesota for example problem 1 16
Figure Soil layers using ldquorules of thumbrdquo17
Figure Soil layer 1 using SBT method 23
Figure Soil layer 2 using SBT method 24
Figure Soil layer 3 using SBT method 25
Figure Soil layer 4 using SBT method 26
Figure CPT data from Minnesota for example problem 2 34
Figure Soil layers using ldquorules of thumbrdquo36
Figure Soil layer 1 using SBT method 42
Figure Soil layer 2 using SBT method 43
Figure Soil layer 3 using SBT method 44
Figure Soil layer 4 using SBT method 45
Figure Dissipation t50 data 53
Figure Direct CPT method for shallow foundations56
Figure Direct CPT method introduction 58
Figure Differentiation of porewater pressure measurement locations (Lunne et al 1997) 58
Figure 25 Foundation soil formation parameter hs versus CPT material index Ic (Mayne 2017) 60
Figure 26 Diagram of footing profiles for Example 362
Figure 27 CPT data from Northern Minnesota 63
Figure 28 Schematic of foundation design associated with CPT data 64
Figure 29 Soil type for example problem 367
Figure 30 Direct CPT evaluation of axial pile capacity 70
Figure 31 UniCone Method soil behavior type using CPT (Mayne 2017) 71
Figure 32 Modified UniCone Method soil behavior type using CPT (Mayne 2017) 71
Figure 33 Deep foundation end bearing pile diagram74
Figure 34 CPT data from Minnesota for example problem 5 75
Figure 35 Soil layers using ldquorules of thumbrdquo for pile capacity example using direct CPT method 76
Figure 36 Soil layer 1 using SBT method 81
Figure 37 Soil layer 2 using SBT method 82
Figure 38 Soil layer 3 using SBT method 83
Figure 39 Soil layering compared to pilings 85
LIST OF TABLES
Table 1 Geoparameters calculated directly from CPT 4
Table 2 Minor Geoparameters 5
Table 3 Soil type compared to exponent m 10
Table 4 Geoparameters evaluated for Case Example 117
Table 5 Geoparameters 35
Table 6 Bearing capacity defined by soil type61
Table 7 SCPT Results 62
CHAPTER 1 INTRODUCTION
The purpose of this manual is to provide an analysis of soil characterization shallow footings and deep
foundations using direct cone penetration testing (CPT) methods Geotechnical site characterization is
important for evaluating soil parameters that will be used in the analysis and design of foundations
retaining walls embankments and situations involving slope stability Common practice is to determine
these soil parameters through conventional lab and in-situ testing An alternative method uses CPT
readings of cone tip resistance (qt) sleeve friction ( fs) and pore pressure (u2) directly to determine these
parameters such as unit weight effective friction angle undrained shear strength and many others
(Figure 1) Direct CPT methods are provided for these parameters which will be used in the designs for
shallow and deep foundations
Figure 1 Geoparameters determined from CPT
Designing shallow foundations is typically done in a two-part process determining the bearing capacity
and expected settlement (commonly referred to as displacement) of the soil to approximate the
required size and shape of a foundation The older traditional methods are no longer required with
many approaches existing for using CPT directly in the design of shallow foundations Results from these
methods can provide a direct assessment of bearing capacity andor settlement A specific approach to
the direct method has been recommended and tested using a database of 166 full-scale field load tests
(Figure 2) A one-part process is used to scale the measured cone penetrometer readings (ie
measured cone tip resistance sleeve friction and pore water pressure) to obtain the bearing capacity
and settlement of the soil A step-by-step procedure has been created to transition from the CPT data to
bearing capacity with settlement accounted for The steps consist of estimating a design footing width
1
and length while using the process in the Soil Characterization section to determine soil parameters
directly from the CPT data
Figure 2 Conventional method for shallow foundation design compared to direct CPT method
There are upwards of 40 different direct CPT methods that have been developed over the past five
decades to determine the axial compression capacity of a piling foundation Earlier direct CPT methods
relied on hand-recorded information where mechanical-type CPT cone tip resistance data would be
collected at 20 cm intervals
Herein the method recommended for deep foundation design is the Modified UniCone method which
uses all three readings of the modern electronic piezocone penetrometer (CPTu) while addressing a
variety of pile foundation types (Figure 3) The modified UniCone Method is based on a total of 330 pile
load tests (three times the original UniCone database) that were associated with SCPTu data Two
computer software programs have been recommended for analyzing pile movements andor
settlements
2
Figure 3 Direct CPT evaluation of axial pile capacity
3
Symbol Parameter
γt Soil total unit weight
Ic CPT material index
SBT Soil behavior type (SBT)
ʹ σp Preconsolidation stress
YSR Yield stress ratio
ϕʹ Effective friction angle
Ko Lateral stress coefficient
su Undrained shear strength
Dʹ Constrained modulus
Eʹ Drained Youngrsquos modulus
CHAPTER 2 DIRECT CPT METHOD FOR SOIL
CHARACTERIZATION
21 INTRODUCTION
A direct CPT method for determining the value of each of various geoparameters shown in Table 1 is
provided The parameter Ic is used in the derivation of several sequential parameters This section of
the guide may be referred to as these parameters are used in calculations throughout the shallow
foundations and deep foundations sections
Table 1 Geoparameters calculated directly from CPT
4
Kʹ Bulk modulus
ks Subgrade reaction modulus
MR Resilient modulus
Gmax Small-strain shear modulus
k Coefficient of permeability
cv Coefficient of consolidation
Symbol Parameter Equation
ρt Mass Density ρt = γtga
where ga = 98 ms2
σvo Total Stress σvo asymp Σ (γti middot Δzi)
c Effective cohesion ʹ Empirical c asymp 003σp
In clays c asymp 01cu
Other minor parameters can be used in calculations of the geoparameters in Table 1 These minor parameters are provided in Table 2
Table 2 Minor Geoparameters
5
22 SOIL UNIT WEIGHT
The total soil unit weight can be estimated from CPT sleeve friction resistance as shown in Figure 4
(Mayne 2014) This method is not applicable to organic clays diatomaceous soils peats or sensitive
soils
Figure 4 Soil unit weight from CPT sleeve friction
Use Equation 1 to calculate the soils total unit weight
1 120574119905 = 120574119908 ∙ [122 + 015 ∙ ln (100 ∙ 119891119904 + 001)]
120590119886119905119898
Since soil unit weight is required for determining most geoparameters (including Soil Behavior Type)
estimating the soil type of the layers using the ldquorules of thumbrdquo method is a good first step before
determining more precise layering (Figure 5) Once a unit weight is determined for each noticeable layer
change these results can be used in later calculations such as ldquoCPT Material Indexrdquo To use the ldquorules
of thumbrdquo method some helpful guidelines are to assume sands are identified when qt gt 725 psi and u2
asymp uo while the presence of intact clays are prevalent when qt lt 725 psi and u2 gt uo The magnitude of
porewater pressures help to indicate intact clays such as soft (u2 asymp 2middotuo) firm (u2 asymp 4middotuo) stiff (u2 asymp 8middotuo)
and hard (u2 asymp 20middotuo) Fissured overconsolidated clays tend to have negative u2 values such that u2 lt 0
6
Figure 5 Approximate ldquorules of thumbrdquo method using CPT sounding from Wakota Bridge MN
23 CPT MATERIAL INDEX
The development of the CPT material index Ic has improved the initial classification of soil types and
calculation of soil parameters as shown in Table 1 To calculate Ic follow steps 1 and 2
231 Step 1 Normalized Sleeve Friction
Calculate Fr using Equation 2 if not provided with the CPT data gathered
2
7
119891119904 119865119903() = 100 ∙
(119902119905minus120590119907119900)
232 Step 2 Iteration
Iterate using Equations 3-5 to determine Ic by initially using n = 1 to calculate a starting value of Ic The
exponent n is soil-type dependent n = 1 (clays) n asymp 075 (silts) and n asymp 05 (sands) Iteration converges
quickly which is generally after the 3rd cycle
3 (119954119957minus120648119959119952)120648119938119957119950 = 119951 119928119957119951 prime (120648119959119952120648119938119957119950)
4
prime 120590119907119900 119899 = 0381 ∙ 119868119888 + 005 ( ) minus 015 120590119886119905119898
119899 le 10
5 119920119940 = radic[120785 120786120789 minus 119949119952119944119928119957119951]120784 + [120783 120784120784 + 119949119952119944119917119955]120784
24 SOIL BEHAVIOR TYPE (SBT)
Typically soil samples are not taken when using CPT The soil types are then inferred from the qt fs and
u2 readings To determine the types of soil from CPT data follow steps 1 and 2
241 Step 1
To determine the soil layers from CPT results calculate Ic by following the steps under section ldquoCPT
Material Indexrdquo After Ic has been determined through all specified depths use Figure 6 to classify the
type of soil by comparing each Ic value to normalized CPT readings (Fr and Qtn) from Equations 2 and 3
8
Figure 6 SBT zones using CPT Ic
242 Step 2
To determine if any of the soil layers contain ldquosensitive clays and siltsrdquo from zone 1 or ldquovery stiff
overconsolidated (OC) soilrdquo from zones 8 and 9 use Equations 6 and 7 If any soil layers are found within
zone 1 by Equation 6 then caution should be taken as these clays are prone to instability collapses and
difficulties in construction performance Very stiff OC sands to clayey sands of zone 8 (15 lt Fr lt 45)
and very stiff OC clays to silts of zone 9 (Fr gt 45) can be identified by Equation 7
6 Equation 6 errata This is an exponential expression (see Figure 5)
119876119905119899 lt 12119890119909119901(minus14 ∙ 119865119903)
7 119876119905119899 gt 0005(119865119903minus1)minus00003(119865119903minus1)2minus0002
1
Errata See terms and coefficients in Figure 5 above (numbers cannot be rounded off)
After each CPT reading has been assigned a zone from Figure 14 a visual representation can be made to
show the predominant layers by soil types (Figure A5)
25 EFFECTIVE STRESS FRICTION ANGLE
The effective friction angle (ϕ) is used to govern the strength for sands and clays where Equation 8 is
used for sands and Equation 9 is used for clays The value of ϕ for sands is derived from Ic so before ϕ
can be calculated refer to the iteration of Qtn under the ldquoCPT Material Indexrdquo section Once the values
9
Soil Type m
Fissured clays 11
Intact clays 10
Sensitive clays 09
Silt mixtures 085
Silty sands 080
Clean sands 072
Note m may be higher than 11 in fissured clays
of Ic and Qtn are calculated the type of soil can be determined from the section ldquoSoil Behavior Type
SBTrdquo The type of soil will dictate which equation to use for ϕ
Sands
8 120601prime(deg) = 176deg + 110deg log(119876119905119899)
Clays
9 119902119905minus120590119907119900 120601prime(deg) = 295deg ∙ 119861119902
0121 ∙ [0256 + 0336 ∙ 119861119902 + log ( )] prime 120590119907119900
Where = 119861119902 (1199062 minus 119906119900)(119902119905 minus 120590119907119900)
26 STRESS HISTORY
Determining the stress history can be characterized by an apparent yield stress ratio of the form
10 prime 120590119901
119884119878119877 = prime 120590119907119900
Where σp is defined as the preconsolidation stress or effective yield stress (Equation 10) The YSR is the
same equation as the more common overconsolidation ratio (OCR) but is now generalized to
accommodate mechanisms of preconsolidation such as ageing desiccation repeated cycles of wetting-
drying repeated freeze thaw cycles and other factors
11 prime prime 120590119901 = 120590119910 = 033(119902119905 minus 120590119907119900)119898prime
Where σp σy σvo and qt have units of kPa and the value of m depends on soil type with typical values shown in Table 3
Table 3 Soil type compared to exponent m
10
The value of m for non-fissured soils and inorganic clays and silts is derived from Ic (Figure 7) so before
m can be calculated (Equation 12) refer to the iteration of Ic under the ldquoCPT Material Indexrdquo section
Determine Ic for all soil layers before calculating m
12 028
119898prime = 1 minus 25 119868119888 1+( frasl ) 265
Once m is known for all soil layers the YSR can be determined
Figure 7 Yield stress exponent compared to CPT material index
A limiting value of YSR can be reached for clays and sands It can be calculated using Equation 13
13 (1 sin(emptyprime))
(1+sin(emptyprime)) 119884119878119877119897119894119898119894119905 = [ ]
(1minussin(emptyprime))2
27 LATERAL STRESS COEFFICIENT
The lateral stress coefficient Ko = σhoσvo commonly referred to as the at-rest condition is used to
represent the horizontal geostatic state of soil stress Ko can be calculated using Equation 14 but
11
14
119870119900 = (1 minus sin(emptyprime)) ∙ 119884119878119877sin( emptyprime)
sections such as ldquoStress Historyrdquo and ldquoEffective Stress Friction Anglerdquo will need to be referred to for
determining parameters ϕ and YSR
A maximum value for Ko can be determined by Equation 15
15 (1+sin(emptyprime))
119870119900119898119886119909 = = 1199051198861198992(45deg + emptyprime2) (1minussin(emptyprime))
28 UNDRAINED SHEAR STRENGTH
Loading on soils can result in fully drained partially drained or fully undrained conditions Sands
typically produce drained cases due to their high permeability but exceptions may occur in loose sands
during fast loading where the water does not have sufficient time to dissipate Clays exhibit low
permeability and thus often result in undrained loading cases when a load is applied quickly For soft-
firm clays the undrained shear strength (su) can be determined from CPT via Equation 16 where the
value of the bearing factor Nkt can be taken as 12
16 119902119905minus120590119907119900 119904119906 =
119873119896119905
In the case of remolded undrained shear strength from CPT su asymp fs
29 GROUND STIFFNESS AND SOIL MODULI
Determining the grounds stiffness can be measured from geoparameters such as the constrained
modulus (D) drained Youngrsquos modulus (E) bulk modulus (K) subgrade reaction modulus (ks) resilient
modulus (MR) and small-strain shear modulus (Gmax)
17 119863 prime asymp 5 ∙ (119902119905 minus 120590119907119900)
18
12
119864 prime = 119863 prime
11
19 119864prime
119870prime = [3∙(1minus2119907prime)]
The subgrade modulus (ks) is a combination of soil-structural properties which creates a parameter that
depends on the ground stiffness and the size of the loaded element
20 119864prime
119896119904 = [119889∙(1minus1199072)]
The resilient modulus MR applies to pavement analysis and design and can be calculated using Equation
21 where qt and fs are in MPa
21 053 119872119877 = (146119902119905 + 135511989111990414 + 236)244
The small strain shear modulus Gmax is a representation of the initial stiffness of all soils and rocks
Graphically it is the beginning portion of all stress-strain-strength curves for geomaterials Use Equation
22 to determine Gmax
22 119866119898119886119909 = 120588119905 ∙ 1198811199042
Where Vs = shear wave velocity (Equation 23) as measured by seismic cone penetration tests (SCPT) If
only cone penetration tests (CPT) or piezocone (CPTu) data are available the shear wave velocity may
be estimated from
23 03 119891119904 119881119904 (119898frasl119904) = [101 ∙ log(119902119905) minus 114]167 ∙ (100 ∙ )
119902119905
Where qt and fs have units of (kPa)
210 COEFFICIENT OF CONSOLIDATION
The coefficient of consolidation (cv) controls the rate that foundation and embankment settlements
occur By using results of CPT dissipation tests that measure the change in u2 readings over time the
value of cv can be determined Using Equation 24 cv can be determined based on piezocone dissipation
13
curves The equation below requires an estimate of the in-situ rigidity index (IR) of the soil If results of
SCPTU are available then IR may be determined from Gmax and qt per equation A38
24 0030∙(119886119888)2∙(119868119877)075
119888119907 = 11990550
Where ac = penetrometer radius (178 cm for 10-cm2 cone 220 cm for 15-cm2 cone) t50 = time to reach 50 dissipation IR = Gsu = undrained rigidity index G can be determined from the ldquoGround Stiffness and Soil Modulirdquo section
211 HYDRAULIC CONDUCTIVITY
The hydraulic conductivity also known as the coefficient of permeability (k) expresses the flow
characteristics of soils and has units of cms or feetday One method of calculation would be to use
Equation 25 where cv would need to be determined from ldquoCoefficient of Consolidationrdquo and D would
need to be determined from ldquoGround Stiffness and Soil Modulirdquo
25 119888119907∙120574119908 119896 =
119863prime
An alternative approach developed for soft normally-consolidated soils (Figure 8) is shown in Equation
26 where t50 (sec) values are used directly in assessing k in (cms)
26
14
125
119896 asymp ( 1
) 251∙11990550
Figure 8 k vs dissipation time for 50 consolidation (Mayne 2017)
212 EXAMPLE PROBLEMS
2121 Example 1 Direct CPT Methods for Geoparameters on Sands
Several geoparameters need to be determined based on the given CPT data collected for the South
Abutment of a bridge in Benton County MN (Figure 9) The groundwater table (GWT) was measured at
17 feet Determine all the geoparameters found in Table 4 at depths of 0 feet to 30 feet All sand layers
can be assumed ldquodrainedrdquo with ν = 02
15
Figure 9 CPT data from Benton County Minnesota for example problem 1
16
Table 4 Geoparameters evaluated for Case Example 1
Symbol Parameter Symbol Parameter
γt Soil total unit weight Ko Lateral stress coefficient
Ic CPT material index Dʹ Constrained modulus
SBT Soil behavior type (SBT) Eʹ Drained Youngrsquos modulus ʹ σp Preconsolidation stress Kʹ Bulk modulus
YSR Yield stress ratio MR Resilient modulus
ϕʹ Effective friction angle Gmax Small-strain shear modulus
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 10 Soil layers using ldquorules of thumbrdquo
γw = 6224 pcf
17
Layer 1
From Figure 10 fs = 17 psi taken as a representative value of the layer
17 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1204 pcf
145 psi
Layer 2
From Figure 10 fs = 17 psi taken as a representative value of the layer
17psi
γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1204 pcf 145 psi
Layer 3
From Figure 10 fs = 12 psi taken as a representative value of the layer
12 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 4
From Figure 10 fs = 7 psi taken as a representative value of the layer
CPT Material Index
7 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1121 pcf
145 psi
Layer 1
From Figure 10 fs = 17 psi and qc = 3500 psi
18
qt = qc + u2 ∙ (1 minus a) = 3500 psi + 3 psi ∙ (1 minus 08) = 3501 psi
= γt ∙ 6 feet = 1204 pcf ∙ 6 feet = 7225 psf = 50 psi σvo
fs 17 psi Fr() = 100 ∙ = 100 ∙ = 049
(qt minus σvo) (3501 psi minus 50 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 5 psi minus 0 psi = 50 psi
(qt minus σvo)σatm (3501 psi minus 50 psi )145 psi = = Qtn prime )n (σvoσatm (50 psi145 psi)n
prime σvo 50 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(049)]2
Qtn = 35329 n = 036 lt 10 Ic = 13
Layer 2
From Figure 10 fs = 17 psi and qc = 3500 psi
19
qt = qc + u2 ∙ (1 minus a) = 3500 psi + 0 psi ∙ (1 minus 08) = 3500 psi
= 7225 psf + γt ∙ 6 feet = 7225 psf + 1204 pcf ∙ 6 feet = 1445 psf = 100 psi σvo
fs 17 psi Fr() = 100 ∙ = 100 ∙ = 049
(qt minus σvo) (3500 psi minus 100 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 100 psi minus 0 psi = 100 psi
(qt minus σvo)σatm (3500 psi minus 100 psi )145 psi = = Qtn prime )n (σvoσatm (100 psi145 psi)n
prime σvo 100 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
47 minus log(Qtn)]2 + [122 + log(049)]2
Ic = radic[3
Qtn = 27947 n = 041 lt 10 Ic = 14
Layer 3
From Figure 10 fs = 12 psi and qc = 1500 psi
20
qt = qc + u2 ∙ (1 minus a) = 1500 psi + 0 psi ∙ (1 minus 08) = 1500 psi
= 1445 psf + γt ∙ 11 feet = 1445 psf + 1172 pcf ∙ 11 feet = 2734 psf = 190 psi σvo
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 081
(qt minus σvo) (1500 psi minus 190 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = γwater ∙ (z minus 119911119908) = 6224 pcf ∙ (23 feet minus 17 feet) = 3734 psf = 99 psi
prime σvo = σvo minus uo = 190 psi minus 99 psi = 164 psi
(qt minus σvo)σatm (1500 psi minus 190 psi )145 psi = = Qtn prime )n (σvoσatm (164 psi145 psi)n
prime σvo 164 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(081)]2
Qtn = 947 n = 06 lt 10 Ic = 19
21
Layer 4
From Figure 10 fs = 7 psi and qc = 1200 psi
qt = qc + u2 ∙ (1 minus a) = 1200 psi + 0 psi ∙ (1 minus 08) = 1200 psi
= 2734 psf + γt ∙ 11 feet = 2734 psf + 1121 pcf ∙ 7 feet = 3519 psf = 244 psi σvo
fs 7 psi Fr() = 100 ∙ = 100 ∙ = 060
(qt minus σvo) (1200 psi minus 244 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = γwater ∙ (z minus 119911119908) = 6224 pcf ∙ (30 feet minus 17 feet) = 8091 psf = 130 psi
prime σvo = σvo minus uo = 244 psi minus 130 psi = 188 psi
(qt minus σvo)σatm (1200 psi minus 244 psi )145 psi = = Qtn prime )n (σvoσatm (188 psi145 psi)n
prime σvo 188 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(060)]2
22
Qtn = 686 n = 064 lt 10 Ic = 19
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic Qtn and Fr the first layer is defined as a ldquoDrained Gravelly Sandrdquo from
Figure 11
Figure 11 Soil layer 1 using SBT method
23
Layer 2
Based on values of Ic Qtn and Fr the second layer is defined as a ldquoDrained Sandrdquo from Figure
12
Figure 12 Soil layer 2 using SBT method
24
Layer 3
Based on values of Ic Qtn and Fr the third layer is defined as a ldquoDrained Sandrdquo from Figure 13
Figure 13 Soil layer 3 using SBT method
25
Layer 4
Based on values of Ic Qtn and Fr the fourth layer is defined as a ldquoDrained Sandrdquo from Figure 14
Figure 14 Soil layer 4 using SBT method
Effective Stress Friction Angle
Layer 1
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(3533) = 456deg
Layer 2
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(2795) = 445deg
26
Layer 3
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(947) = 393deg
Layer 4
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(686) = 378deg
Stress History
Layer 1
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (13frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(24139 kPa minus 345 kPa)072 = 4717 kPa
= 684 psi
prime σp 684 psi YSR = = = 136
primeσvo 50 psi
Layer 2
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + (14 1 + ( frasl ) frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(34132 kPa minus 689 kPa)072 = 605 kPa
= 877 psi
prime σp 877 psi YSR = = = 87
primeσvo 100 psi
27
Layer 3
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + (19 1 + ( frasl ) frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(10342 kPa minus 131 kPa)072 = 254 kPa
= 368 psi
prime σp 368 psi YSR = = = 22
primeσvo 164 psi
Layer 4
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (19frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(8274 kPa minus 168 kPa)072 = 215 kPa = 312 psi
prime σp 312 psi YSR = = = 17
primeσvo 188 psi
Lateral Stress Coefficient
Layer 1
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(456deg)) ∙ 136sin( 456deg) = 18
Layer 2
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(445deg)) ∙ 87sin( 445deg) = 14
28
Dprime 17480 psi
Eprime = = = 15890 psi 11 11
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(393deg)) ∙ 22sin( 393deg) = 06
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(378deg)) ∙ 17sin( 378deg) = 05
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3501 psi minus 50 psi) = 17480 psi
Eprime 15890 psi
Kprime = = = 8828 psi [3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Layer 3
Layer 4
Ground Stiffness and Soil Moduli
Layer 1
Values of qt and fs need to be in MPa
MR = (146(241 MPa)053 + 1355(012 MPa)14 + 236)244 = 3416 MPa = 49575 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1172 kPa m Vs (mfrasls) = [101 ∙ log(24139 kPa) minus 114]167 ∙ (100 ∙ ) = 2745
24139 kPa s
Vs = 9012 fts
29
γt 1204 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 37 ∙ (9012 ) = 304 ∙ 106psf = 21000 psi s
Layer 2
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3500 psi minus 100 psi) = 17480 psi
Dprime 17480 psi Eprime = = = 15890 psi
11 11
Eprime 15890 psi Kprime = = = 8828 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(241 MPa)053 + 1355(012 MPa)14 + 236)244 = 3415 MPa = 49562 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1172 kPa m Vs (mfrasls) = [101 ∙ log(24133 kPa) minus 114]167 ∙ (100 ∙ ) = 2745
24133 kPa s
Vs = 9012 fts
γt 1204 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
30
2 ft Gmax = ρt ∙ Vs
2 = 37 ∙ (9012 ) = 304 ∙ 106psf = 21000psi s
Layer 3
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (1500 psi minus 190 psi) = 7405 psi
Dprime 7405 psi Eprime = = = 6732 psi
11 11
Eprime 6732 psi Kprime = = = 3740 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(103 MPa)053 + 1355(008 MPa)14 + 236)244 = 1506 MPa = 21854 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 827 kPa m Vs (mfrasls) = [101 ∙ log(10343 kPa) minus 114]167 ∙ (100 ∙ ) = 2611
10343 kPa s
Vs = 8566 fts
γt 1172 pcf slug ρt = = = 36
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 36 ∙ (8566 ) = 27 ∙ 106psf = 19000 psi s
Layer 4
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (1200 psi minus 244 psi) = 5878 psi
31
Dprime 5878 psi Eprime = = = 5343 psi
11 11
Eprime 5343 psi Kprime = = = 2969 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(83 MPa)053 + 1355(005 MPa)14 + 236)244 = 165 MPa = 16901 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 483 kPa m Vs (mfrasls) = [101 ∙ log(8274 kPa) minus 114]167 ∙ (100 ∙ ) = 2243
8274 kPa s
Vs = 7360 fts
γt 1121 pcf slug ρt = = = 35
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 35 ∙ (7360 ) = 389 ∙ 106psf = 13000 psi s
32
2122 Example 2 Direct CPT Methods for Geoparameters on Clay
Several geoparameters need to be determined based on the given CPT data collected for the South
Abutment (Figure 15) The groundwater table (GWT) was measured at 60 feet Determine all the
geoparameters found in Table 5 at depths of 0 feet to 42 feet All sand layers can be assumed ldquodrainedrdquo
ν = 02 All clay layers can assume ν = 049
33
Figure 15 CPT data from Minnesota for example problem 2
34
Table 5 Geoparameters
Symbol Parameter Symbol Parameter
γt Soil total unit weight Eʹ Drained Youngrsquos modulus
Ic CPT material index Kʹ Bulk modulus
SBT Soil behavior type (SBT) MR Resilient modulus
ʹ σp Preconsolidation stress Gmax Small-strain shear modulus
YSR Yield stress ratio su Undrained shear strength
ϕʹ Effective friction angle cv Coefficient of consolidation
Ko Lateral stress coefficient k Hydraulic conductivity
Dʹ Constrained modulus
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
35
Figure 16 Soil layers using ldquorules of thumbrdquo
Unit weight of water γw = 6224 pcf
Layer 1
From Figure 16 fs = 13 psi taken as a representative value of the layer
13 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1179 pcf
145 psi
Layer 2
From Figure 16 fs = 12 psi taken as a representative value of the layer
12 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 3
From Figure 16 fs = 20 psi taken as a representative value of the layer
36
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
Layer 4
From Figure 16 fs = 2 psi taken as a representative value of the layer
2 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1004 pcf
145 psi
CPT Material Index
Layer 1
From Figure 10 fs = 13 psi and qc = 3000 psi
qt = qc + u2 ∙ (1 minus a) = 3000 psi + 0 psi ∙ (1 minus 08) = 3000 psi
σvo = γt ∙ 2 feet = 1179 pcf ∙ 2 feet = 235 psf = 16 psi
fs 13 psi Fr() = 100 ∙ = 100 ∙ = 043
(qt minus σvo) (3000 psi minus 16 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 16 psi minus 0 psi = 16 psi
(qt minus σvo)σatm (3000 psi minus 16 psi )145 psi = = Qtn prime )n (16 psi145 psi)n (σvoσatm
prime σvo 16 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(043)]2
37
Qtn = 4126 n = 032 lt 10 Ic = 121 therefore sand
Layer 2
From Figure 10 fs = 12 psi and qc = 250 psi
qt = qc + u2 ∙ (1 minus a) = 250 psi + 10 psi ∙ (1 minus 08) = 252 psi
σvo = 235 psf + γt ∙ 30 feet = 235 psf + 1172 pcf ∙ 30 feet = 3751 psf = 260 psi
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 53
(q minus σ ) (252 psi minus 260 psi) t vo
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 260 psi minus 0 psi = 260 psi
(qt minus σvo)σatm (250 psi minus 260 psi )145 psi = = Qtn prime (σvoσatm)n (260 psi145 psi)n
38
prime σvo 260 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(53)]2
Qtn = 87 n = 10 le 10 Ic = 32 therefore clay
Layer 3
From Figure 10 fs = 20 psi and qc = 4000 psi
qt = qc + u2 ∙ (1 minus a) = 4000 psi + 8 psi ∙ (1 minus 08) = 40016 psi
σvo = 3751 psf + γt ∙ 2 feet = 3751 psf + 1219 pcf ∙ 2 feet = 3995 psf = 277 psi
fs 16 psi Fr() = 100 ∙ = 100 ∙ = 054
(qt minus σvo) (30016 psi minus 277 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 277 psi minus 0 psi = 277 psi
(qt minus σvo)σatm (30016 minus 277)145 psi = = Qtn prime (σvoσatm)n (277 psi145 psi)n
39
primeσvo 277 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(054)]2
Qtn = 1962 n = 052 le 10 Ic = 15 therefore sand
Layer 4
From Figure 10 fs = 2 psi and qc = 250 psi
qt = qc + u2 ∙ (1 minus a) = 250 psi + 0 psi ∙ (1 minus 08) = 250 psi
= 3994 psf + γt ∙ 8 feet = 3994 psf + 1004 pcf ∙ 8 feet = 4798 psf = 333 psi σvo
fs 2 psi Fr() = 100 ∙ = 100 ∙ = 092
(qt minus σvo) (250 psi minus 333 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 333 psi minus 0 psi = 333 psi
(qt minus σvo)σatm (250 psi minus 333 psi )145 psi = = Qtn prime (σvoσatm)n (333 psi145 psi)n
40
prime σvo 333 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(092)]2
Qtn = 65 n = 10 lt 10 Ic = 29 therefore clayey silt
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic Qtn and Fr the first layer is defined as a ldquoDrained Sandrdquo from Figure 17
41
Figure 17 Soil layer 1 using SBT method
Layer 2
Based on values of Ic Qtn and Fr the second layer is defined as a ldquoUndrained Clayrdquo from Figure
18
42
Figure 18 Soil layer 2 using SBT method
43
Layer 3
Based on values of Ic Qtn and Fr the third layer is defined as a ldquoDrained Sandrdquo from Figure 19
Figure 19 Soil layer 3 using SBT method
44
Layer 4
Based on values of Ic Qtn and Fr the fourth layer is defined as a ldquoUndrained Silty Mixrdquo from
Figure 20
Figure 20 Soil layer 4 using SBT method
Effective Stress Friction Angle
Layer 1 (sand)
45
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(4126) = 464deg
Layer 2 (clay)
qt minus σvo
ϕprime(deg) = 295deg ∙ Bq0121 ∙ [0256 + 0336 ∙ Bq + log ( )]
primeσvo
Where
(u2 minus uo) (10 psi minus 0 psi) Bq = = = 004
(qt minus σvo) (252 psi minus 260 psi)
252 psi minus 260 psi ϕprime(deg) = 295deg ∙ 00440121 ∙ [0256 + 0336 ∙ 0044 + log ( )]
260 psi
ϕprime(deg) = 245deg
Layer 3 (sand)
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(1962) = 428deg
Layer 4 (clay)
qt minus σvo
ϕprime(deg) = 295deg ∙ Bq0121 ∙ [0256 + 0336 ∙ Bq + log ( )]
primeσvo
Where
(u2 minus uo) (5 psi minus 0 psi) Bq = = = 002
(qt minus σvo) (251 psi minus 333 psi)
252 psi minus 333 psi ϕprime(deg) = 295deg ∙ 0020121 ∙ [0256 + 0336 ∙ 002 + log ( )]
333 psi
ϕprime(deg) = 265deg
Stress History
46
Layer 1
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (12frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(3000 psi minus 33)072 = 1051 psi
prime σp 1051 psi YSR = = = 642
primeσvo 16 psi
Layer 2
028 028
prime m = 1 minus = 1 minus = 099 25 25 1 + (
Icfrasl ) 1 + (32frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(252 psi minus 260)099 = 734 psi
prime σp 734 psi YSR = = = 28
primeσvo 260 psi
Layer 3
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + ( frasl ) 1 + (15frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(4002 psi minus 277)072 = 1288 psi
47
prime σp 1288 psi YSR = = = 46
primeσvo 277 psi
Layer 4
028 028
prime m = 1 minus = 1 minus = 097 25 25 Ic 1 + ( frasl ) 1 + (29frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(250 psi minus 333)097 = 626 psi
prime σp 626 psi YSR = = = 19
primeσvo 333 psi
Lateral Stress Coefficient
Layer 1
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(464deg)) ∙ 642sin( 464deg) = 56
Layer 2
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(245deg)) ∙ 28sin( 245deg) = 090
Layer 3
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(428deg)) ∙ 46sin( 428deg) = 091
Layer 4
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(266deg)) ∙ 19sin( 266deg) = 073
48
Ground Stiffness and Soil Moduli
Layer 1
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3000 psi minus 16 psi) = 14991 psi
Dprime 14991 psi Eprime = = = 13629 psi
11 11
Eprime 13629 psi Kprime = = = 7571 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(207 MPa)053 + 1355(009 MPa)14 + 236)244 = 2816 MPa = 40873 psi
fs
03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 896 kPa m Vs (mfrasls) = [101 ∙ log(20685 kPa) minus 114]167 ∙ (100 ∙ ) = 2564
20685 kPa s
Vs = 8412 fts
γt 1179 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
2 ft 2 Gmax = ρt ∙ Vs = 37 ∙ (8412 ) = 260 ∙ 106psf = 17992 psi
s
Layer 2
49
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (252 psi minus 260 psi) = 11298 psi
Dprime 11298 psi Eprime = = = 10271 psi
11 11
Eprime 10271 psi Kprime = = = 17118 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(049))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(17 MPa)053 + 1355(008 MPa)14 + 236)244 = 443 MPa = 64309 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 827 kPa m Vs (mfrasls) = [101 ∙ log(17375 kPa) minus 114]167 ∙ (100 ∙ ) = 2646
17375 kPa s
Vs = 8680 fts
γt 1172 pcf slug ρt = = = 36
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 36 ∙ (8680 ) = 274 ∙ 106psf = 19036 psi s
Layer 3
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (4002 psi minus 277 psi) = 19869 psi
Dprime 19869 psi Eprime = = = 18063 psi
11 11
50
Eprime 18063 psi Kprime = = = 10035 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(276 MPa)053 + 1355(014 MPa)14 + 236)244 = 4017 MPa = 58309 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1379 kPa m Vs (mfrasls) = [101 ∙ log(27591 kPa) minus 1379]167 ∙ (100 ∙ ) = 2854
27591 kPa s
Vs = 9363 fts
γt 121 pcf slug ρt = = = 38
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 38 ∙ (9363 ) = 33 ∙ 106psf = 23050 psi s
Layer 4
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (250 psi minus 333 psi) = 1088 psi
Dprime 1088 psi Eprime = = = 989 psi
11 11
Eprime 989 psi Kprime = = = 16490 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(049))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
51
MR = (146(17 MPa)053 + 1355(001 MPa)14 + 236)244 = 360 MPa = 52189 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 138 kPa m Vs (mfrasls) = [101 ∙ log(17238 kPa) minus 114]167 ∙ (100 ∙ ) = 1545
17238 kPa s
Vs = 5069 fts
γt 1004 pcf slug ρt = = = 31
ga 322 ftfrasls2 ft3
2 ft 2 Gmax = ρt ∙ Vs = 31 ∙ (5069 ) = 80 ∙ 105psf = 5565 psi
s
Undrained Shear Strength
Layer 2
qt minus σv0 250 psi minus 26 psi su = = = 187 psi
12 12
Layer 4
qt minus σv0 250 psi minus 333 psi su = = = 181 psi
12 12
Coefficient of Consolidation
52
Example dissipation data are shown in Figure 21 Example calculations are provided
Figure 21 Dissipation t50 data
Layer 1
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (7197)075
cv = = = 359 cmsec 1 sec t50
Layer 2
53
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (94392)075
cv = = = 0008 cmsec 3000 sec t50
Layer 3
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (8303)075
cv = = = 665 cmsec 06 sec t50
Layer 4
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (26714)075
cv = = = 087 cmsec 11 sec t50
Hydraulic Conductivity
Layer 1
cm 1 psi 359 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 10 ∙ 10minus4 cmsec Dprime 14984 psi
Layer 2
cm 1 psi 0008 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 73 ∙ 10minus7 cmsec Dprime 11379 psi
Layer 3
cm 1 psi 665 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 20 ∙ 10minus5 cmsec Dprime 148650 psi
54
Layer 4
cm 1 psi 087 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 18 ∙ 10minus4 cmsec Dprime 10861 psi
55
CHAPTER 3 DIRECT CPT METHOD FOR SHALLOW
FOUNDATIONS
31 PROCEDURE
Shallow foundation analysis is typically done in a two-part traditional procedure The traditional
techniques are no longer required as a direct CPT method for square rectangular and circular shallow
footings is available (Figure 22) This process has the soil types grouped into four main categories sands
silts fissured clays and intact clays When determining soil types for each design it is believed footings
on sands and silts act in a fully drained manner while intact clays act in an undrained manner under
conditions of constant volume In order to determine the vertical stress-displacement-capacity of
square rectangular and circular shallow footings follow the steps provided towards the solution given
by Equation 27
Figure 22 Direct CPT method for shallow foundations
Equation 27 may be used to calculate all footing stresses from zero to the bearing capacity (qmax) To
calculate qmax for a sized footing of width (B) length (L) and thickness (t) follow steps 1 through 7
provided The settlement (s) can be determined after the calculation of qmax by simply rearranging
56
Equation 27 where the allowable stress (qallow) is defined as qmax divided by the factor of safety (FS) For
shallow footings a FS value of 3 is commonly used in geotechnical engineering
120782120787 minus120782120785120786120787 119956 119923 119954119950119938119961 = 119945119956 ∙ 119954119957119951119942119957 ∙ (119913
) ∙ (119913
) 27 119950119938119961
2 119902119898119886119909 0345 1 frasl119865119878 119871 119904 = 119861 ∙ [ ∙ ∙ ( ) ] 28
119861 ℎ119904 119902119905119899119890119905
Where hs = the foundation soil formation parameter qtnet = the net corrected cone tip resistance
311 Step 1 Estimating Footing Dimensions
In Minnesota frost heave can have devastating effect on a shallow foundations It is common practice to
place a foundation bearing elevation below the expected maximum frost depth (roughly 45 to 6 feet
below ground level) With this assumption estimate a footing size (B x L) for design to obtain
representative data from CPT roughly 15middotB below the foundation depth (Df) as shown in Figure 23 The
CPT data collected will consist of
cone tip resistance (qc)
measured porewater pressure acting behind the cone tip (u2) as shown in Figure 24
sleeve friction (fs)
57
Figure 23 Direct CPT method introduction
Figure 24 Differentiation of porewater pressure measurement locations (Lunne et al 1997)
312 Step 2 Soil Characterization
The soil behavior type (SBT) and CPT material index (Ic) govern the value of the formation factor hs The
first step is following the steps in Soil Unit Weight to determine soil layering and γt values for each
layer To determine a representative unit weight (γsoil) for all the layers in the range of Df to 15∙B below
58
Df the user will need to use their engineering judgement on the definition of ldquorepresentative unit
weightrdquo This single value of unit weight is used in further calculations such as the total vertical soil stress (σvo) from Equation 29 and effective vertical stress (σvo) from Equation 30 Values of σvo and σvo
will need to be calculated at 15middotB below Df
120648119959119952 = sum(120632119956119952119946119949 ∙ 119963) 29
prime 120648119959119952 = 120648119959119952 minus 119958119952 30
Where z = Df +15middotB uo=γwatermiddot(z-zw)
The qc will need to be corrected using Equation 31 in the case of fine-grained soils that develop excess porewater pressure during cone penetration These values will be used in the following calculations
119954119957 = 119954119940 + 119958120784 ∙ (120783 minus 119938) 31
Where 119860119899 a = cone area ratio = eg MnDOT commonly uses a = 08 119860119888
119860119899 = cross-sectional area of load cell or shaft 119860119888 = projected area of the cone
The cone area ratio is determined based on the type of piezocone tip used during in-situ field testing
Manufacturer specifications should provide the measured net area ratio (a) for the particular cone
penetrometer as determined by calibration in a pressurized triaxial chamber
313 Step 2a Foundation soil formation parameter
With the representative value γsoil continue to follow the steps in CPT Material Index through Soil
Behavior Type (SBT) to better define the type of soil at depth 15∙B below Df Use the value of Ic
calculated at depth 15∙B below Df to calculate hs with Equation 36 This parameter is based on the soil
type with typical values shown in Figure 25 The data on silts and sands are considered fully drained
whereas the fissured clay subset may be partially drained to undrained
59
23 ℎ119904 = 28 minus
119868119888 15 32
1+( ) 24
Figure 25 Foundation soil formation parameter hs versus CPT material index Ic (Mayne 2017)
314 Step 3 Soil elastic modulus and Poissonrsquos ratio
Details about the soil such as its elastic modulus (Es) and Poissonrsquos ratio (ν) will be needed for further
calculations A representative value for the Es can be determined from in-situ field tests Values of ν can
be assigned as 02 for drained sands and as 05 for undrained loading cases involving clays (Jardine et al
1985 Burland 1989)
315 Step 4 Net cone tip resistance
Calculate qtnet the mean value of net cone tip resistance 15middotB below the foundation bearing elevation
using Equation 33
33 119902119905119899119890119905 = 119902119905 minus 120590119907119900
60
316 Step 5 Bearing capacity of the soil
Use the assumed B and L calculated hs and qtnet to determine the soils bearing capacity (qmax) from
Equation 34 Use Table 6 to calculate qmax by using the maximum allowable settlement ratio (sB)max
correlating to the soil type If hs is in between the given values interpolate to acquire (sB)max
Table 6 Bearing capacity defined by soil type
Type of Soil hs (sB)max
Clean Sands 058 12
Silts 112 10
Fissured Clays 147 7
Intact Clays 270 4
05 minus0345 119904 119871 119902119898119886119909 = ℎ119904 ∙ 119902119905119899119890119905 ∙ (119861
) ∙ (119861
) 34 119898119886119909
317 Step 6 Settlement
Settlement can be calculated directly using the results from Equation 35 A FS equal to 3 is common in
foundation engineering
2 119902119898119886119909 0345 1 frasl119865119878 119871 119904 = 119861 ∙ [ ∙ ∙ ( ) ] 35
119861 ℎ119904 119902119905119899119890119905
318 Step 7 Final Check
Check that the applied stress (q) is less than qmax using Equation 36 Repeat the process again with a
new B andor new L if q gt qmax
05 119904 ) 119902 = ℎ119904 ∙ 119902119905119899119890119905 ∙ (119861
minus0345 119871 ∙ ( )
11986136
61
32 EXAMPLE PROBLEMS
321 Example 3 Direct CPT Method on Sands
A footing size needs to be determined based on the given CPT data collected for the South Abutment
(Figure 27) The footing stress (q) was determined to be 8000 psf Estimate a footing size (B x L) and
determine the bearing capacity of the foundation (qmax) using the direct CPT method provided Also
determine the expected settlement based on the calculated bearing capacity
Figure 26 Diagram of footing profiles for Example 3
The soil elastic modulus determined from seismic CPT (SCPT) are shown in Table 7
Table 7 SCPT Results
Depth (feet) Es (tsf)
Bottom of layer
3 557
6 433
9 557
12 695
17 590
22 501
27 571
32 505
62
Figure 27 CPT data from Northern Minnesota
Solution
Step 1 Assume the footing is placed below the frost depth of 6 feet
Estimate L L = 50 feet = 600 inches
63
Estimate B B = 12 feet = 144 inches
Estimate footing thickness t t = 2 feet
Df = 6 feet = 72 inches
Df + 15middotB = 24 feet = 288 inches
Step 2 Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 28 Schematic of foundation design associated with CPT data
Layer 1
From Figure 28 fs = 20 psi taken as a representative value of the layer
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
64
Layer 2
From Figure 28 fs = 20 psi taken as a representative value of the layer
20psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
Layer 3
From Figure 28 fs = 8 psi taken as a representative value of the layer
8 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1134 pcf
145 psi
Using the unit weights calculated between Df and 15B determine a representative unit weight of the soil to calculate the total and effective soil stresses Based on layer 3 being the weakest supporting layer γsoil = 113 pcf
σvo = γsoil ∙ (119863119891 + 15 ∙ B) = 113 lbsfraslft3 ∙ 24 feet = 2721 psf = 189 psi
prime σvo = σvo minus uo = 2721 psf minus 0 psf = 2721 psf = 189 psi
Calculate the cone tip resistance between Df and Df + 15B below the foundation depth
qt = qc + u2 ∙ (1 minus a) = 1250 psi + 0 psi ∙ (1 minus 08) = 1000 psi
Step 2a CPT Material Index
Using Figure 28 the sleeve friction at 15B below the foundation depth is about 16 psi
65
fs 16 psi Fr() = 100 ∙ = 100 ∙ = 13
(qt minus σvo) (1250 psi minus 189 psi)
Iterate to solve for Ic Steps are not shown for brevity
(qt minus σvo)σatm (1250 psi minus 189 psi )145 psi = = Qtn prime (σvoσatm)n (189 psi145 psi)n
prime σvo 189 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Qtn = 703 n = 072 lt 10 Ic = 210
66
Figure 29 Soil type for example problem 3
Use the Ic value to determine the foundation soil formation parameter
Calculate the foundation soil formation parameter The tip resistance is about 1250 psi at 24 feet and
the cone area ratio was determined to be 08 from information provided by the manufacturer
23 23 hs = 28 minus = 28 minus = 078 = SandSilt 15 15 Ic 210
1 + ( ) 1 + ( ) 24 24
67
Step 3 Determine representative values of soil elastic modulus from soil testing and Poissons ratio The
soil elastic modulus was taken as the average value between depths of Df and 15middotB (6 feet and 24 feet)
433 + 557 + 695 + 590 + 501 Es = = 555 tsf = 708 psi
5
ldquoDrained sandsiltrdquo gives a ν = 020
Step 4 Calculate the net cone tip resistance
qtnet = qt minus σvo = 1250 psi minus 189 psi = 12311 psi
Step 5 Calculate the bearing capacity of the sand
In drained sandssilts the bearing capacity is taken as the stress when (sB) = 011 (or 11 foundation width)
minus0345 minus0345 05 s L 600 qmax = hs ∙ qtnet ∙ ( ) ∙ ( ) = 078 ∙ (9795) ∙ (011)05 ∙ ( )
B B 144
= 193 psi = 27860 psf qmax
Assuming a factor of safety (FS) of 3
qmax = 645 psi = 9287 psf
3
Step 6 Calculate settlement
68
119902119898119886119909 0345 2
1 frasl119865119878 L 119904 = 119861 ∙ [ ∙ ∙ ( ) ]
ℎ119904 119902119905119899119890119905 B
2 0345 1 645 psi 600 inches s = 144 inches ∙ [ ∙ ∙ ( ) ] = 18 inches
078 12311 psi 144 inches
Step 7 Determine if q gt qmax
q = 8000 psf lt 9287psf = qmax3
69
CHAPTER 4 DIRECT CPT METHOD FOR DEEP FOUNDATIONS
41 INTRODUCTION
The axial compression capacity (Qtotal) for a single pile includes a side component (Qside) end bearing
component (Qbase) and pile weight (Wpile) as shown in Figure 30 Since piles commonly push through
several layers a summation of the unit side frictions acting on the pile segments must be considered
over the length of the pile While the examples shown in this Guide resemble hand calculations
computer software is more efficient Programming the procedure is possible and represents a practical
method of designing deep foundations However commercial software is frequently available and is
MnDOTrsquos most common method of designing deep foundations
There are upwards of 40 different direct CPT methods that have been developed over the past five
decades to determine a piles axial compression capacity Many of the earliest methods relied on hand-
recorded information where qc data from mechanical CPTs would be collected at 20 cm intervals
whereas the direct CPT method uses scaled penetrometer readings via specified algorithms to obtain
the pile unit side friction and unit end bearing The method that will be used for deep foundation design
is the Modified UniCone method which uses all three readings of the electronic piezocone (qt fs and u2)
while addressing a variety of pile foundation types
Figure 30 Direct CPT evaluation of axial pile capacity
70
The modified UniCone Method is based upon a total of 330 pile load tests (three times the original
Unicone database) that were associated with SCPTu data Originally the UniCone method provided
approximate soil classification in five groups via a chart of qE vs fs (Figure 31) where qE = qt -u2 Later
using the modified approach with a larger data set provided soil sub classifications as shown in Figure
32 This new 9-zone normalized soil behavior type is determined using CPT data in combination with the
CPT Material Index
Figure 31 UniCone Method soil behavior type using CPT (Mayne 2017)
Figure 32 Modified UniCone Method soil behavior type using CPT (Mayne 2017)
71
42 MODIFIED UNICONE METHOD
In order to determine the axial pile capacity using the modified method the first step requires the
determination of geoparameters as shown in Direct CPT Method for Soil Characterization Once the
soil unit weight and CPT Material index are determined for each soil layer then the effective cone
resistance pile unit side friction (fp) and pile end bearing resistance (qb) can be determined
421 Step 1
Work through the steps provided in CPT Method for Soil Characterization until each soil layer is
defined by its CPT material index and soil behavior type using Figure 6 After these steps have been
completed continue to Step 2 to determine qE qb and fp
422 Step 2
Once qt is determined for each layer the effective cone resistance can be calculated using Equation 37
119902119864 = 119902119905 minus 1199062 37
Where (a) qE is the specific value at each elevation along the pile sides for determining fp and (b) at the bottom of the pile qE is averaged in the vicinity of the pile tip from the tip bearing elevation to about one diameter beneath the tip for determining qb
Using CPT material index and qE the pile end bearing resistance is calculated using Equation 38
38 119902119887 = 119902119864 ∙ 10(0325∙119868119888minus1218)
The pile unit side friction is obtained from qE and Ic
39 119891119901 = 119902119864 ∙ 120579119875119879 ∙ 120579119879119862 ∙ 120579119877119860119879119864 ∙ 10(0732∙119868119888minus3605)
Where θPT = coefficient for pile type (084 for bored 102 for jacked 113 for driven piles) θTC = coefficient for loading direction (111 for compression and 085 for tension) θRATE = rate coefficient applied to soils in SBT zone 1 through 7 (109 for constant rate of penetration test and 097 for maintained load tests)
72
423 Step 3
Due to piles extending through multiple layers the unit side components acting on various pile
segments would need to be summed if not using the direct CPT method Since CPT calculates data at
regular intervals of 2 cm to 5 cm along the sides of the pile the average fp in each layer can be used
directly in Equation 40 to obtain the shaft capacity
119876119904119894119889119890 = 119891119901 ∙ 119860119904 40
Where fp = average pile side friction along pile length from eqn 39 As = πmiddotdmiddotH d = pile diameter and H = length embedded below grade
The base capacity for a pile in compression loading is given by Equation 41 For piles in tension (or uplift)
Qbase can be taken as 0
= 119902119887 ∙ 119860119887 41 119876119887119886119904119890 Where
qb = end bearing resistance from eqn 38 Ab = πmiddotd24 (area of a circular pile)
424 Step 4
The final step is to calculate the axial pile capacity using Equation 42
119876119905119900119905119886119897 = 119876119904119894119889119890 + 119876119887119886119904119890 minus 119882119901119894119897119890 42
43 AXIAL PILE DISPLACEMENTS
Movement of pile foundations can be assessed using elastic continuum theory which has been
developed using finite element analyses boundary elements and analytical closed-form solutions In
the case of piles passing through several soil layers the elastic solution can be used by stacking pile
segments (each with its own stiffness) as represented by soil Youngrsquos modulus The use of software is
recommended for pile groups Several available programs such as DEFPIG GROUP and PIGLET can
handle pile groups under axial and lateralmoment loading
73
44 EXAMPLE PROBLEMS
441 Example 5 Direct CPT Methods Axial Pile Capacity
Several piles need to be placed beneath the edge of a building Use the CPT data collected for this site
shown in Figure 34 to determine the axial capacity for one of the piles Assume round steel driven piles
will be used with a diameter of 1275 inches wall thickness of 025 inches and lengths of 80 feet
Concrete will be used to fill the piles To solve for all geoparameters use the Direct CPT Method for Soil
Characterization section All sand layers can be assumed ldquodrainedrdquo with ν = 02
Figure 33 Deep foundation end bearing pile diagram
74
Figure 34 CPT data from Minnesota for example problem 5
75
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 35 Soil layers using ldquorules of thumbrdquo for pile capacity example using direct CPT method
Layer 1
From Figure 35 fs = 18 psi taken as a representative value of the layer
76
18 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1210 pcf
145 psi
Layer 2
From Figure 35 fs = 12 psi taken as a representative value of the layer
12psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 3
From Figure 35 fs = 20 psi taken as a representative value of the layer
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
CPT Material Index
Layer 1
From Figure 35 fs = 18 psi and qc = 3000 psi
qt = qc + u2 ∙ (1 minus a) = 3000 psi + 20 psi ∙ (1 minus 08) = 3004 psi
∙ 4 feet = 1219 pcf ∙ 4 feet = 4877 psf = 34 psi σvo = γt
fs 18 psi Fr() = 100 ∙ = 100 ∙ = 060
(qt minus σvo) (3004 psi minus 34 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
77
uo = 0 psi
prime σvo = σvo minus uo = 34 psi minus 0 psi = 34 psi
(qt minus σvo)σatm (3004 psi minus 34 psi )145 psi = = Qtn prime )n (σvoσatm (34 psi145 psi)n
prime σvo 34 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(060)]2
Qtn = 3593 n = 038 lt 10 Ic = 14 (ie sand)
Layer 2
From Figure 35 fs = 12 psi and qc = 500 psi
qt = qc + u2 ∙ (1 minus a) = 500 psi + 40 psi ∙ (1 minus 08) = 508 psi
= 4838 psf + γt ∙ 45 feet = 4838 psf + 1172 pcf ∙ 45 feet = 5756 psf = 400 psi σvo
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 260
(qt minus σvo) (508 psi minus 400 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
78
uo = 0 psi
prime σvo = σvo minus uo = 400 psi minus 0 psi = 400 psi
(qt minus σvo)σatm (508 psi minus 400 psi )145 psi = = Qtn prime (σvoσatm)n (400 psi145 psi)n
prime σvo 400 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(260)]2
Qtn = 117 n = 10 le 10 Ic = 29 (ie clayey silt)
Layer 3
From Figure 35 fs = 20 psi and qc = 5000 psi
qt = qc + u2 ∙ (1 minus a) = 5000 psi + 0 psi ∙ (1 minus 08) = 5000 psi
= 5756 psf + γt ∙ 6 feet = 5756 psf + 1219 pcf ∙ 6 feet = 6488 psf = 451 psi σvo
fs 20 psi Fr() = 100 ∙ = 100 ∙ = 040
(qt minus σvo) (5000 psi minus 451 psi)
79
Step 2 Iterate to solve for Ic Steps are not shown for brevity
prime σvo = σvo minus uo = 451 psi minus 0 psi = 451 psi
(qt minus σvo)σatm (5000 psi minus 451 psi )145 psi = = Qtn prime )n (σvoσatm (451 psi145 psi)n
prime σvo 451 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(040)]2
= 1801 n = 06 lt 10 = 15 (ie sand) Qtn Ic
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic (14) Qtn (359) and Fr (06) the first layer is defined as a ldquoGravelly Sandrdquo
from Figure 36
80
Figure 36 Soil layer 1 using SBT method
81
Layer 2
Based on values of Ic (29) Qtn (117) and Fr (26) the second layer is defined as a ldquoSilty Mixrdquo
from Figure 37
Figure 37 Soil layer 2 using SBT method
82
Layer 3
Based on values of Ic (15) Qtn (180) and Fr (04) the third layer is defined as a ldquoSandrdquo from
Figure 38
Figure 38 Soil layer 3 using SBT method
Effective Cone Resistance
Layer 1
83
qE = qt minus u2 = 3004 psi minus 20 psi = 2984 psi
Layer 2
qE = qt minus u2 = 508 psi minus 40 psi = 468 psi
Layer 3
qE = qt minus u2 = 5000 psi minus 0 psi = 5000 psi
Pile End Bearing Resistance
Layer 3
= qE ∙ 10(0325∙Icminus1218) qb = 5000 psi ∙ 10(0325∙15minus1218) = 9085 psi
Pile Unit Side Friction
Layer 1
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 2984 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙14minus3605) = 99 psi
Layer 2
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 468 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙29minus3605) = 211 psi
Layer 3
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 5000 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙15minus3605) = 202 psi
84
Axial Pile Capacity
The side capacity is based on pile design as shown in Figure 39
Layer 1
Qside = fp ∙ As = 99 psi ∙ (π ∙ d ∙ L) = 99 psi ∙ (π ∙ 1275 in ∙ 48 in) = 19064 lb
Layer 2
Qside = fp ∙ As = 211 psi ∙ (π ∙ d ∙ L) = 211 psi ∙ (π ∙ 1275 in ∙ 588 in) = 457122 lb
Layer 3
Qside = fp ∙ As = 202 psi ∙ (π ∙ d ∙ L) = 202 psi ∙ (π ∙ 1275 in ∙ 240 in) = 58170 lb
Figure 39 Soil layering compared to pilings
85
End Bearing Resistance in Layer 3
d2 1275 in2
Qbase = qb ∙ Ab = 9085 psi ∙ (π ∙ ) = 9085 psi ∙ (π ∙ ) = 115932 lb 4 4
Wp = (γsteel ∙ Asteel ∙ Depth) + (γsteel ∙ Aconc ∙ Depth)
Wp = (490 pcf ∙ 003 ft2 ∙ 55 ft) + (150 pcf ∙ 085 ft2 ∙ 55 ft) = 7724 lbs
Layer 3
Qtotal = ΣQside + Qblayer3 minus Wp
Qtotal = (19064 + 45713 + 58170) lbs + 115932 lbs minus 7724 lbs = 642563 lbs
= 642 kips Qtotal
Table 8 Summary of selected and calculated parameters
Layer L (in) qc
(psi)
fs
(psi)
Fr Ic Qtn qE
(psi)
qb
(psi)
Qbase
(lb)
fp
(psi)
Qside
(lb)
Qtotal
(kip)
1 48 3000 18 06 14 356 2984 NA NA 99 19064
2 588 500 12 26 29 117 468 NA NA 211 457122
3 240 5000 20 04 15 180 5000 9085 115932 202 58170
Total 115932 534355 642
86
REFERENCES
Abu-Farsakh MY Zhang Z Tumay M and Morvant M (2008) Computerized cone penetration test for soil classification Transportation Research Record 2053 47-64
Agaiby SA (2018) Advancements in the interpretation of seismic piezocone tests in clays and other geomaterials PhD dissertation School of Civil amp Environmental Engineering Georgia Institute of Technology Atlanta GA
Agaiby SS and Mayne PW (2018) Interpretation of piezocone penetration and dissipation tests in sensitive Leda Clay at Gloucester Test Site Canadian Geotechnical Journal (accepted for publication 08 March 2018) CGJ-2017-0388
Amar S Baguelin F Canepa Y and Frank R (1998) New design rules for the bearing capacity of shallow foundations based on Menard PMT Geotechnical Site Characterization Vol 2 (Proc ISC-1 Atlanta) 727-733
American Petroleum Institute (API) (1981) Planning designing and constructing fixed offshore platforms Recommended practice API-RP2A 17th edition Washington DC API
American Petroleum Institute (API) (1987) Planning designing and constructing fixed offshore platforms Recommended practice API-RP2A 18th edition Washington DC API
American Petroleum Institute (API) (1989) Planning designing and constructing fixed offshore platforms Recommended practice API-RP2A 19th edition Washington DC API
Andersen KH and Stenhamar P (1982) Static plate loading tests on overconsolidated clay Journal of the Geotechnical Engineering Division ASCE 108 (GT7) 919-934
Basu D and Salgado R (2012) Load and resistance factor design of drilled shafts in sands Journal of Geotechnical amp Geoenvironmental Engineering 138 (12) 1455-1469
Berardi R Jamiolkowski M and Lancellotta R (1991) Settlement of shallow foundations in sands selection of stiffness on the basis of penetration resistance Geotechnical Engineering Congress Vol 1 185-200 (Proc GSP-27 Boulder) Reston Virginia ASCE
Brand EW Muktabhant C and Taechathummarak A (1972) Load tests on small foundations in soft clay Performance of Earth and Earth-Supported Structures Vol 1 Part 2 903-928 (Proc PC) ASCE Reston Virginia ASCE
Briaud JL and Gibbens RM (1999) Behavior of five large spread footings in sand Journal of Geotechnical amp Geoenvironmental Engineering 125 (9) 787-797
Briaud JL (2007) Spread footings in sand load-settlement curve approach Journal of Geotechnical amp Geoenvironmental Engineering 133 (8) 905-920
Brown DA Turner JP and Castelli RJ (2010) Drilled Shafts Construction Procedures and LRFD Design Methods Geotechnical Engineering Circular GEC 10 Report No FHWA NHI-10-016 Washington DC National Highway Institute US DOT
87
Burland J B (1973) Shaft Friction of Piles in Clay -A Simple Fundamental Approach Ground Eng 6(3) 30-42
Cai G Liu S and Puppala A (2012) Reliability assessment of CPTu-based pile capacity predictions in soft clay deposits Engineering Geology 84-91
Canadian Geotechnical Society (1992) Canadian Foundation Engineering Manual Third Edition Vancouver British Columbia BiTech Publishers
Cerato AB and Lutenegger AJ (2007) Scale effects of shallow foundation bearing capacity on granular material Journal of Geotechnical amp Geoenvironmental Engineering 133 (10) 1192-1201
Chen BS-Y and Mayne PW (1996) Statistical relationships between piezocone measurements and stress history of clays Canadian Geotechnical Journal 33 (3) 488-498
Cho W and Finno RJ (2010_ Stress-strain responses of block samples of compressible Chicago glacial clays Journal of Geotechnical amp Geoenvironmental Engineering 136 (1) 178-188
Chung SF Randolph MF and Schneider JA (2006) Effect of penetration rate on penetrometer resistance in clay Journal of Geotechnical amp Geoenvironmental Engineering 132 (9) 1188-1196
Clausen CJF Aas PM and Karlsrud K (2005) Bearing capacity of driven piles in sand the NGI approach Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis Group
Consoli NC Schnaid F and Milititsky J (1998) Interpretation of plate load tests on residual soil site Journal of Geotecnical amp Geoenvironmental Engingeering 124 (9) 857-867
Dasenbrock DD (2006) Assessment of pile capacity by static and dynamic methods to reconcile design predictions with observed performance Proc 54th Minnesota Geotechnical Engineering Conference 95-108 edited by Labuz J Univ of Minnesota St Paul
DeBeer E and Martens A (1957) Method of computation of an upper limit for the influence of heterogeneity of sand layers on the settlement of bridges 275-282 Proc ICSMFE-4 Vol 1 London ICSMFE
DeBeer EE (1965) Bearing capacity and settlement of shallow foundations on sand Proc Symposium on Bearing Capacity and Settlement of Foundations 15-33 Duke University Durham NC
Decourt L (1999) Behavior of foundations under working load conditions Proc PCSMGE-11 Foz do Iguassu Brazil Vol 4 453-487
Demers D and Leroueil S (2002) Evaluation of preconsolidation pressure and the overconsolidation ratio from piezocone tests of clay deposits in Quebec Canadian Geotechnical Journal 39 (1) 174-192
Eslaamizaad S and Robertson PK (1996) Cone penetration test to evaluate bearing capacity of foundations in sand 429-438 Proc CGCFG-49 Vol 1 St Johns Newfoundland
Eslami A Alimirzaei M Aflaki E and Molaabasi H (2017) Deltaic soil behavior classification using CPTu records Marine Georesources amp Geotechnology 35 (1) 62-79
88
Eslami A and Gholami M (2005) Bearing capacity analysis of shallow foundations from CPT data 1463-1466 Proc ICSMGE- 16 Vol 3 (Osaka) Millpress Rotterdam
Eslami A and Gholami M (2006) Analytical model for the ultimate bearing capacity of foundations from cone resistance Scientia Iranica 13 (3) 223-233
Eslami A and Fellenius BH (1997) Pile capacity by direct CPT and CPTu methods applied to 102 case histories Canadian Geotechnical Journal 34 (6) 886-904
Fellenius BH (2009) Basics of Foundation Design Vancouver BiTech Publishers
Fellenius BH and Altaee A (1994) Stress and settlement of footings in sand Vertical and Horizontal Deformations of Foundations and Embankments 2 1760-1773
Fellenius BH and Eslami A (2000) Soil profile interpreted from CPTu data Proc Geotechnical Engineering Conference Year 2000 Geotechnics Asian Inst of Technology Bangkok Thailand Available from wwwfelleniusnet
Finno R J and Chung C K (1992) Stress-strain-strength responses of compressible Chicago glacial clays Journal of Geotechnical Engineering 118 (10) 1607ndash1625
Fleming K Weltman A Randolph M and Elson K (2009) Piling Engineering Third Edition London Taylor amp Francis Group
Frank R and Magnan J-P (1995) Cone penetration testing in France national report Proceedings International Symposium on Cone Penetration Testing 3 (CPT95 Linkoumlping) Swedish Geotechnical Society Report 3(95) 147-156
Gavin K Adekunte A and OKelly B (2009) A field investigation of vertical footing response on sand Geotechnical Engineering 162 257-267
Hegazy YA (1998) Delineating geostratigraphy by cluster analysis of piezocone data PhD dissertation School of Civil amp Environmental Engineering Georgia Institute of Technology Atlanta GA
Holtz RD Kovacs WD and Sheehan TC (2011) An Introduction to Geotechnical Engineering 2nd Edition London Pearson Publishing
Hunt CE Pestana JM Bray JD and Riemer M (2002) Effect of pile driving on static and dynamic properties of soft clay Journal of Geotechnical amp Geoenvironmental Engineering 128 (1) 13-24
Jamiolkowski M Ladd CC Germaine JT and Lancellotta R (1985) New developments in field and lab testing of soils Proc ICSMFE-11 1 57-154
Jardine R Chow F Overy R and Standing J (2005) ICP design methods for driven piles in sands and clays London Thomas Telford Publishing
Jardine RJ Lehane BM Smith P and Gildea P (1995) Vertical loading experiments on rigid pad foundations at Bothkennar Geotechnique 45 (4) 573-597
89
Karlsrud K (2012) Prediction of load-displacement behavior and capacity of axially-loaded piles in clay based on analyses and interpretation of pile load test results PhD Dissertation Norwegian Univ Science amp Technology Trondheim Norway
Karlsrud K Clausen CJF and Aas PM (2005) Bearing capacity of driven piles in clay the NGI approach Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis
Karlsrud K Lunne T Kort DA and Strandvik S (2001) CPTU correlations for clays Proc 16th ICSMGE 2 683-702
Kimura T Kusakabe O and Saitoh K (1985) Geotechnical model tests of bearing capacity problems in a centrifuge Geotechnique 35 (1) 33-45
Knappett JA and Craig RF (2012) Craigs Soil Mechanics 8th Edition London Spon Press Taylor amp Francis
Kolk HJ and Velde EVD (1996) A reliable method to determine friction capacity of piles driven into clays Paper No 7993 Proc Offshore Technology Conference Houston
Kolk HJ Baaijens AE and Senders M (2005) Design criteria for pipe piles in silica sands Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis Group
Kulhawy FH (2004) On the axial behavior of drilled shafts Geosupport 2004 1-18
Kulhawy FH and Jackson CS (1989) Foundation Engineering Current Principles amp Practices 2 (GSP 22) Reston VA ASCE
Kulhawy FH and Mayne PW (1990) Manual for Estimating Soil Properties for Foundation Design EPRI Report EL-6800 Electric Power Research Institute Palo Alto 306 page Download from wwwepricom
Kulhawy FH Trautmann CH Beech JF ORourke TD and McGuire W (1983) Transmission Line Structure Foundations for Uplift-Compression Loading Report EL-2870 Palo Alto CA Electric Power Research Institute wwwepricom
Ku T and Mayne PW (2015) In-situ lateral stress coefficient (K0) from shear wave velocity measurements in soils Journal of Geotechnical amp Geoenvironmental Engineering 141 (12) 101061(ASCE) GT1943-56060001354
Ladd CC and DeGroot DJ (2003) Recommended practice for soft ground site characterization Soil amp Rock America 2003 3-57
Larsson R (1997) Investigations and Load Tests in Silty Soils SGI Report R-54 Linkoumlping Swedish Geotechnical Institute
Larsson R (2001) Investigations and Load Tests in Clay Till SGI Report R-59 Linkoumlping Swedish Geotechnical Institute
Lee J and Salgado R (2005) Estimation of bearing capacity of circular footings on sands based on cone penetration tests Journal of Geotechnical amp Geoenvironmental Engineering 131 (4) 442-452
90
Lehane BM (2003) Vertically loaded shallow foundation on soft clayey silt Geotechnical Engineering 156 (1) Thomas Telford London 17-26
Lehane BM (2013) Relating foundation capacity in sands to CPT qc Geotechnical amp Geophysical Site Characterization 1 London Taylor amp Francis Group
Lehane BM Doherty JP and Schneider JA (2008) Settlement prediction for footings on sand Deformational Characteristics of Geomaterials (IS-Atlanta) 1 Rotterdam Millpress
Lehane BM Li Y and Williams R (2013) Shaft capacity of displacement piles in clay using the CPT Journal of Geotechnical amp Geoenvironmental Engineering 139 (2) 253-266
Lehane BM Schneider JA and Xu X (2005) The UWA-05 method for prediction of axial capacity of driven piles in sand Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis
Low HE Lunne T Andersen KH Sjursen MA Li X and Randolph MF 2010 Estimation of intact and remoulded undrained shear strengths from penetration tests in soft clays Geotechnique 60 (11) 843-859
Lunne T Robertson PK and Powell JJM (1997) Cone Penetration Testing in Geotechnical Practice London RoutledgeBlackie Academic
Lutenegger AJ and DeGroot DJ (1995) Settlement of Shallow Foundations on Granular Soils Report Contract 6332 Task Order 4 submitted to Massachusetts Highway Dept by University of Mass-Amherst Amherst MA
Lutenegger AJ and Adams MT (2003) Characteristic load-settlement behavior of shallow foundations Proc FONDSUP 2 381-392
Mase T and Hashiguchi K (2009) Numerical analysis of footing settlement problem by subloading surface model Soils amp Foundations 49 (2) 207-220
Mayne PW (1987) Determining preconsolidation stress and penetration pore pressures from DMT contact pressures Geotechnical Testing Journal 10(3) 146-150
Mayne PW (1991) Determination of OCR in Clays by Piezocone Tests Using Cavity Expansion and Critical State Concepts Soils and Foundations 31 (2) 65-76
Mayne PW (2007) NCHRP Synthesis 368 on Cone Penetration Test Washington DC Transportation Research Board National Academies Press wwwtrborg
Mayne PW (2009) Geoengineering Design Using the Cone Penetration Test Richmond BC Canada ConeTec Inc
Mayne PW (2014) Keynote Interpretation of geotechnical parameters from seismic piezocone tests Proceedings 3rd Intl Symp on Cone Penetration Testing 47-73
Mayne PW (2016) Evaluating effective stress parameters and undrained shear strengths of soft-firm clays from CPT and DMT Australian Geomechanics Journal 51 (4) 27-55
91
Mayne PW (2017) Stress history of soils from cone penetration tests 34th Manual Rocha Lecture Soils amp Rocks Vol 40 (3) Saouml Paulo ABMS
Mayne PW Christopher B Berg R and DeJong J (2002) Subsurface Investigations -Geotechnical Site Characterization Publication No FHWA-NHI-01-031 Washington DC National Highway Institute Federal Highway Administration
Mayne PW Coop MR Springman S Huang A-B and Zornberg J (2009) Geomaterial Behavior and Testing Proc 17th Intl Conf Soil Mechanics amp Geotechnical Engineering 4 Rotterdam MillpressIOS Press
Mayne PW and Illingworth F (2010) Direct CPT method for footings on sand using a database approach Proc ISCPT-2 3 315-322
Mayne PW and Niazi F (2017) Recent developments and applications in field investigations for deep foundations Proceedings 3rd Bolivian Conference on Deep Foundations 1 141-160
Mayne PW Peuchen J and Baltoukas D (2015) Piezocone evaluation of undrained strength in soft to firm offshore clays Frontiers in Offshore Geotechnics III 2 1091-1096
Mayne PW and Poulos HG (1999) Approximate displacement influence factors for elastic shallow foundation systems ASCE Journal of Geotechnical amp Geoenvironmental Engineering 125 (6) 453-460
Mayne PW Uzielli M and Illingworth F (2012) Shallow footing response on sands using a direct method based on cone penetration tests Full Scale Testing and Foundation Design (GSP 227) Reston Virginia ASCE
Mayne PW and Woeller DJ (2014) Direct CPT method for footings on sands silts and clays Geocharacterization for Modeling and Sustainability (GSP 234) 1983-1997
McClelland B (1974) Design of Deep Penetration Piles for Ocean Structures Journal Geot Eng Div 100(GT7) 705-747
Mesri G (1975) Discussion on new design procedure for stability of soft clays ASCE Journal of the Geotechnical Engineering Division 101(GT4) pp 409-412
Meyerhof GG (1956) Penetration tests and bearing capacity of cohesionless soils Journal of Soil Mechanics amp Foundations Division 82 (SM1) 1-19
Meyerhof GG (1974) General Report Penetration testing outside Europe Proceedings ESOPT-1 Vol 21 Swedish Geotechnical Society Stockholm
Meyerhof GG (1976) Bearing capacity and settlement of pile foundations Journal of Geotechnical Engineering Division 102(3) 197-228
Missiaen T Verhegge J Heirman K and Crobe P (2015) Potential of cone penetrating testing for mapping deeply buried palaeolandscapes in the context of archaeological surveys in polder areas Journal of Archaeological Science 55 174-187
92
Mlynarek Z Wierzbicki J Goglik S and Bogucki M (2014) Shear strength and deformation parameters of peat and gyttja from CPTu DMT and VST Proceedings 5th International Workshop on CPTu and DMT in soft clays and organic soils 193-210 Poznan Poland Polish Committee on Geotechnics Menard Polska Tensar Internationa
Nejaim PF Jannuzzi GMF and Danziger FAB (2016) Soil behavior type of the Sarapuiacute II test site Geotechnical amp Geophysical Site Characterization 5 1009-1014
Niazi FS and Mayne PW (2013) Cone penetration test based direct methods for evaluating static axial capacity of single piles Geotechnical and Geological Engineering 31 (4) 979-1009
Niazi FS and Mayne PW (2015) Enhanced Unicone expressions for axial pile capacity evaluation from piezocone tests Proc International Foundations Conference amp Equipment Expo RestonVA ASCE
Niazi FS and Mayne PW (2016) CPTu-based enhanced UniCone method for pile capacity Engineering Geology 212 21-34
Nowacki F Karlsrud K and Sparrevik P (1996) Comparison of recent tests on OC clay and implications for design Large Scale Pile Tests in Clay London Thomas Telford
ONeill MW (2001) Side resistance in piles and drilled shafts Journal of Geotechnical amp Geoenvironmental Engineering 127 (1) 3-16
Ouyang Z and Mayne PW (2018) Effective friction angle of clays and silts from piezocone Canadian Geotechnical Journal in press doiorg101139cgj-2017-0451
Paikowsky SG Canniff MC Lesny K Kisse A Amatya S and Muganga R (2010) LRFD Design and Construction of Shallow Foundations for Highway Bridge Structures NCHRP Report 651 Washington DC National Cooperative Highway Research Program Transportation Research Board
Pestana JM Hunt CE and Bray JD (2002) Soil deformation and excess pore pressure filed around a closed-ended pile Journal of Geotechnical amp Geoenvironmental Engineering 128 (1) 1-12
Poulos HG and Davis EH (1974) Elastic Solutions for Soil and Rock Mechanics New York Wiley amp Sons
Poulos HG and Davis E (1980) Pile Foundation Analysis amp Design New York John Wiley amp Sons
Powell JJM and Lunne T (2005) Use of CPTU data in clays and fine-grained soils Studia Geotechnica et Mechanica XXVII(3) 29-66
Randolph MF (2003) Rankine Lecture Science and empiricism in pile foundation design Geotechnique 53 (10) 847-875
Robertson PK (1990 1991) Soil classification using the cone penetration test Canadian Geotechnical Journal 27 (1) 151-158 Closure 28 (1) 176-178
Robertson PK (1991) Estimation of foundation settlements in sand from CPT Geotechnical Engineering Congress 2 Reston Virginia ASCE
93
Robertson PK (2009) Cone penetration testing a unified approach Canadian Geotechnical J 46 (11) 1337-1355
Robertson PK (2009a) Performance based earthquake design using the CPT Keynote Lecture International Conference on Performance-Based Design in Earthquake Geotechnical Engineering IS Tokyo Tsukuba Japan
Robertson PK (2009b) Interpretation of cone penetration tests a unified approach Canadian Geotechnical Journal 46 (11) 1337-1355
Robertson PK and Cabal KL (2007) Guide to cone penetration testing Second Edition Signal Hill California Gregg In-Situ
Robertson PK and Cabal KL (2015) Guide to Cone Penetration Testing for Geotechnical Engineering 6th Edition Signal Hill California Gregg Drilling amp Testing Inc
Schmertmann JH (1970) Static cone to compute static settlement over sand Journal of the Soil Mechanics and Foundations Division 95 (SM3) 1011-1043
Schmertmann JH (1978) Guidelines for cone penetration test performance and design Report FHWA-TS-78-209 Washington DC Federal Highway Administration
Schnaid F Wood WR Smith AKC and Jubb P (1993) Investigation of bearing capacity and settlements of soft clay deposits at Shellhaven Predictive Soil Mechanics London Thomas Telford
Schneider JA Xu X and Lehane BM (2008) Database assessment of CPT-based design methods for axial capacity of driven piles in siliceous sands J Geotechnical and Geoenvironmental Engineering 134 (9) 1227-1244
Semple D (1980) Recent Developments in the Design amp Construction of Piles London Institution of Civil Engineers
Senneset K Sandven R amp Janbu N (1989) Evaluation of soil parameters from piezocone tests In-Situ Testing of Soil Properties for Transportation Transportation Research Record 1235 24-37
Shahri AA Melehmir A and Juhlin C (2015) Soil classification analysis based on piezocone penetration test data Engineering Geology 189 32-47
Stuedlein AW and Holtz RD (2010) Undrained displacement behavior of spread footings in clay The Art of Foundation Engineering Practice GSP 198 Reston Virginia ASCE
Takesue K Sasao H and Matsumoto T (1998) Correlation between ultimate pile skin friction and CPT data Geotechnical Site Characterization 2 1177ndash1182
Tand KE Funegard EG and Briaud J-L (1986) Bearing capacity of footings on clay CPT method Use of In-Situ Tests in Geotechnical Engineering GSP 6 1017-1033
Tand KE Warden PE and Funegaringrd EG (1995) Predicted-measured bearing capacity of shallow footings on sand PISCPT 2 589-594
Thomas D (1968) Deep soundings test results and the settlement of spread footings on normally consolidated sands Geotechnique 18 (2) 472-488
94
Tomlinson MJ (1957) The adhesion of piles driven into clay soils Proc 4th Intl Conf on Soil Mechanics and Foundation Engineering 2 66-71
Tomlinson MJ (1986) Foundation Design amp Construction London Longman Scientific
Tomlinson MJ (1987) Pile Design and Construction Practice London Viewpoint Publication
Trofimenkov JG (1974) Penetration testing in the USSR Proceedings ESOPT-1 1 147-154
Tumay MT Hatipkarasulu Y Marx ER and Cotton B (2013) CPTPCPT-based organic material profiling Proc 18th Intl Conf on Soil Mechanics amp Geotechnical Engineering Paris 633-636
Uzielli M and Mayne PW (2011) Statistical characterization and stochastic simulation of load-displacement behavior of shallow footings Proc Georisk 2011 Geotechnical Risk Management amp Assessment (GSP 224) 672-679
Uzielli M and Mayne PW (2012) Load-displacement uncertainty of vertically-loaded shallow footings on sands and effects on probabilistic settlement estimation Georisk Assessment and Management of Risk for Engineered Systems and GeoHazards 6 (1) 50-69
Valsson SM (2016) Detecting quick clay with CPTu Proceedings 17th Nordic Geotechnical Meeting Reykajavik Iceland Challenges in Nordic Geotechnics
Van Dijk BFJ and Kolk HJ (2011) CPT-based design method for axial capacity of offshore piles in clay Frontiers in Offshore Geotechnics II 555-560
Vesić AS (1975) Bearing capacity of shallow foundations Foundation Engineering Handbook New York Van Nostrand Reinhold Publishers
Vesic AS (1977) NCHRP Synthesis of Highway Practice 42 Design of Pile Foundations Washington DC Transportation Research Board
Yu F and Yang J (2012) Base capacity of open-ended steel pipe piles in sand J Geotechnical and Geoenvironmental Engineering 138 (9) 1116-1128
Zawrzykraj P Rydelek P and Bakowska A (2017) Geoengineering properties of Eemian peats from Radsymin (central Poland) in the light of static cone penetration and dilatometer tests Engineering Geology 226 290-300
95
APPENDIX A
List of Figures
Figure A1 Conventional methods versus direct CPT approach to shallow foundation response A-3
Figure A2 Direct bearing capacity relationship for embedded square and strip footings situated on clean
sands (after Schmertmann 1978) A-6
Figure A3 Direct CPT bearing factor Rk as function of footing width B and embedment depth from finite
element analyses (Tand et al 1995) A-6
Figure A4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio and sand
consistency (Eslaamizaad amp Robertson 1996) A-7
Figure A5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp Salgado (2005) in
terms of footing size relative density and base settlement A-8
Figure A6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and normalized
cone tip resistance (after Eslami amp Gholami 2005) A-8
Figure A7 Direct relationship between ultimate bearing stress in clays and measured cone tip resistance
(Schmertmann 1978) A-10
Figure A8 Direct relationship between ultimate foundation bearing stress in clays and net cone tip
resistance (after Tand et al 1986) A-11
Figure A9 Measured load-displacements for each of the five large footings at TAMU (Briaud amp Gibbens
1994) sand site A-12
Figure A10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU (Briaud amp
Gibbens 1994) footings A-13
Figure A11 Characteristic stress versus square root normalized displacements for TAMU footings A-15
Figure A12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sand A-16
Figure A13 Summary of sand formation factors with corresponding CPT resistances A-19
Figure A14 Normalized foundation stresses vs square root of normalized displacement for spread
footings on sand (modified after Mayne et al 2012) A-20
Figure A15 Definitions of capacity from three types of observed foundation load test responses
(modified after Kulhawy 2004) A-21
Figure A16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footings A-22
A-1
Figure A17 Measured load vs displacement curves for 3 shallow foundations on clay at Baytown Texas
(Stuedlein amp Holtz 2010) A-25
Figure A18 Characteristic stress vs square root of normalized displacement for all three foundations at
Baytown site (Stuedlein amp Holtz 2010) A-26
Figure A19 Normalized foundation stress vs square root of normalized displacements A-27
Figure A20 Applied foundation stress vs CPT-calculated stress for 67 footings A-28
Figure A21 Applied footing stress vs CPT calculated stress on logarithmic scales A-29
Figure A22 Summary graph and equations of direct CPT method for evaluating the vertical A-30
Figure A23 Foundation soil formation parameter hs versus CPT material index Ic A-32
Figure A24 Bearing capacity ratio qmaxqnet versus CPT material index Ic A-33
Figure A25 Influence factor for rectangular foundations from elastic theory solution (Mayne amp
Dasenbrock 2017) A-34
List of Tables
Table A1 Direct CPT methods for bearing capacity of footings on clean sands A-4
Table A2 Direct CPT methods for bearing capacity of footings on clays A-9
Table A3 Summary of large footings soil conditions and reference sources of database A-17
Table A4 Sources for shallow foundation performace A-34
A-2
A1 INTERPRETATION OF SOIL PARAMETERS FROM CONE
PENETRATION TESTS
A11 Introduction
Geotechnical site characterization is an important first step towards the evaluation of
subsurface conditions and determination of soil layering geomaterial classification and the
evaluation of soil engineering parameters for the analysis and design of foundations retaining
walls tunnels excavations embankments and slope stability Increasing use of the electronic
cone penetrometer in highway applications is prominent in the USA because the results are
obtained much faster and less expensive than traditional methods that rely on rotary drilling
augering sampling and laboratory testing Moreover the cone penetration test (CPT) provides
at least three independent and continuous readings on soil behavior that are digitally recorded
and fully available at the end of the sounding As such the reliable and consistent
interpretation of CPT data is important since civil engineering works will best realize the
efficiency and economy of this technology in constructed facilities for county state and
interstate projects
A111 Cone Penetration Testing
The cone penetrometer is an electronically-instrumented steel probe that is vertically pushed into the
ground at constant rate of 20 mms (08 insec) using a hydraulic system The penetrometer is outfitted
with load cells pressure transducers and inclinometers that measure at least three readings
continuously with depth (a) cone tip resistance qt (b) sleeve friction fs and (c) porewater pressure u2
With the latter reading the device is called a piezocone thus the designation CPTu is used to indicate
porewater pressure Additional sensors are available that can be used to measure inclination
resistivity pH temperature shear wave velocity and other readings if desired
Figure A1 shows a schematic rendering of the cone penetration test that is performed in the
field per ASTM D 5778 procedures The hydraulic system is nominally rated at 180 kN (20 tons)
capacity and often mounted on a truck but can also be positioned on tracked vehicles or
anchored frames Figure A2 shows a MnDOT cone truck that uses the full dead weight of the
rig for the hydraulic reaction forces which are applied at the center of the vehicle The cab is
enclosed so that soundings may proceed during inclement weather and the data acquisition
system and operator are protected
Figure A 1 Setup and procedures for co ne penetration testing (CPT)
A-4
Figure A2 CPT truck-mounted rig used by MnDOT
A representative sounding taken at the I-35 test site just northeast of Saint Paul is presented in Figure
A3 Here the separate profiles of qt fs and u2 with depth are shown in side-by-side plots In the last
graph the hydrostatic porewater pressure (uo) due to the groundwater table is indicated by the blue
dashed line
These readings are used to interpret the soil layers types and properties of the ground as
discussed in the following sections
A-5
Figure A3 Representative CPT sounding at I-35 test site northeast of St Paul MN showing profiles of
(a) cone resistance (b) sleeve friction and (c) penetration porewater pressures
A112 Soil Parameter Interpretation
A variety of soil engineering parameters can be interpreted from CPT results on the basis of
theoretical frameworks analytical models or numerical simulations otherwise by empirical
methods based on correlations and statistics of databases Figure A4 shows a conceptual
utilization of the readings from the CPT for interpretation of geostratigraphy compressibility
flow characteristics and stress-strain-strength behavior of soils
Table A1 lists a selection of geoparameters that have been addressed for these purposes The
most common or useful relationships will be discussed in subsequent sections and reference is
made to other sources (Lunne et al 1997 Mayne 2007 Robertson amp Cabal 2015) for additional
details
A-6
Figure A4 Conceptual post-processing of CPT results for geoparameter evaluation
Table A1 List of common geoparameters interpreted from CPT data
Symbol Parameter Remarks Notes
SBT Soil behavior type (SBT)
Ic CPT material index
γt Unit weight
ρt Mass Density
σvo Overburden stress
u0 Hydrostatic pressure
A-7
σvo Effective stress
Table A1 Continued
Symbol Parameter Remarks Notes
σp Preconsolidation stress σp = 033 (qnet)m where m = fctn (Ic)
(or Effective Yield Stress) where stresses are in kPa
YSR Yield stress ratio YSR = σpσvo (formerly OCR)
c Effective cohesion intercept
ϕ Effective friction angle
su Undrained shear strength
D Constrained modulus
G0 Small-strain modulus
G Shear modulus
E Youngs modulus
cvh Coefficient of consolidation
K Bulk modulus
A-8
LI
K0 Lateral stress coefficient
k Hydraulic conductivity
Liquefaction index
A12 Geostratigraphy and Soil Behavioral Type (SBT)
In routine CPT soil samples are not normally taken and thus the measured stress friction andor
porewater pressure readings are used to infer the types of soils This can be accomplished using
three basic approaches
a Rules of Thumb or approximate guidelines for a quick visual assessment
b Soil Behavioral Type (SBT) Charts that are based on either the three readings (qt fs u2) net readings
including net cone resistance (qnet = (qt-σvo) effective cone resistance (qE = qt - u2) friction ratio (Rf =
100middotfsqt) and excess porewater pressures or using normalized CPT readings such as Q F Bq or U
c Probabilistic methods where the CPT readings have been calibrated from lab tests on
recovered soil samples (eg Tumay et al 2013)
A121 Approximate Rules of Thumb
The approximate rules of thumb provide simple guidelines for soil type and suggest that sands are
identified when qt gt 5 MPa and u2 asymp u0 whereas intact clays are prevalent when qt lt 5 MPa and u2 gt u0
(Mayne et al 2002) The magnitude of porewater pressures reflects the consistency of the intact clay
such that soft (u2 asymp 2middotu0) firm (u2 asymp 4middotu0) stiff (u2 asymp 8middotu0) and hard (u2 asymp 20middotu0) However for fissured
overconsolidated clays the measured porewater pressures are less than hydrostatic in fact often
negative u2 lt 0 On land negative porewater pressure readings at the shoulder filter element of the
piezocone should be greater than -1 bar ie u2 gt -100 kPa (-147 psi) Also for clean sands the friction
ratio (FR = 100middotfsqt lt 1) while for insensitive clays (FR gt 4)
A-9
Figure A5 presents a 36-m deep piezocone sounding from the MnDOT Wakota Bridge site which crosses
the Mississippi River at I-494 (Dasenbrock 2006) The qt and u2 readings have been annotated using the
aforementioned rules of thumb indicating the predominance of sands at this site As noted there are
five interbedded clay layers evident at depths of 1 m 3-4 m 7 - 12 m 17 m and 22-27 m
Figure A5 Interpretation of soil types from CPT data taken from the Wakota Bridge on I-494 using
approximate rules of thumb
A122 Soil Behavioral Type Charts
The most common approach for identification of soil types is based on soil behavioral charts
Reviews of several chart methods are provided elsewhere (Kulhawy amp Mayne 1990 Fellenius amp
Eslami 2000 Shahri et al 2015) Current popularity favors the 9-zone classification system to
evaluate soil behavioral type (SBT) that uses normalized piezocone readings (Robertson 1990
1991 Lunne et al 1997)
A-10
(119902119905 minus 120590119907119900) (a) normalized tip resistance 119876 = A11 prime 120590119907119900
100∙119891119904 (b) normalized sleeve friction 119865() = A12
(119902119905 minus 120590119907119900)
(1199062 minus 1199060) (c) normalized porewater pressure 119861119902 = A13
(119902119905 minus 120590119907119900)
While the Q-F-Bq classification is a three-dimensional plot it is often presented as a set of two-dimensional
plots specifically Q vs F and Q vs Bq as shown in Figure A6 Data are grouped according to layers and
can be superimposed to identify their association with the nine soil behavioral types All indirect CPT soil
classification approaches should be cross-checked and verified for a particular geologic setting and local
geotechnical conditions before routine use in practice
Figure A6 Nine-zone soil behavioral type charts (a) normalized cone resistance vs normalized
sleeve friction (b) normalized cone resistance vs normalized porewater pressure parameters
(adapted after Robertson 1991 Lunne et al 1997)
The development of a CPT material index (Ic) has been found advantageous in the initial screening of soil
types and helps to organize the sounding into their respective SBT zones of similar soil response In this
case the CPT index is found from (Robertson 2009)
A-11
A2 22 )log221()log473( rtnc FQI
where Qtn = modfied stress-normalized net cone resistance given by
(119902119905 minus 120590119907119900)120590119886119905119898 119876 = A31 prime (120590119907119900120590119886119905119898)119899
where σatm = reference stress equal to atmospheric pressure (1 atm asymp 1 bar = 100 kPa asymp 1 tsf asymp 147 psi)
The exponent n is soil-type dependent n = 1 (clays) n asymp 075 (silts) and n asymp 05 (sands) If units of bars are
used then Qtn (units of bars) is simply
(119902119905 minus 120590119907119900) 119876 = 119899 A32
prime ) (120590119907119900
The operational value of exponent n is found by iteration Initially an exponent n = 1 is used to calculate
the starting value of Ic (ie Qtn = Q) and then the exponent is upgraded to
prime 120590119907119900 119899 = 0381 ∙ 119868119888 + 005 ( ) minus 015 le 10 A4 120590119886119905119898
Then the index Ic is recalculated Iteration converges quickly and generally only 3 cycles are needed to
secure the operational Ic at each depth The soil zones and associated Ic values are detailed in Figure A7
The sensitive soils of zone 1 can be screened using the following expression
A-12
Zone 1 119876119905119899 lt 12119890119909119901( minus14 ∙ 119865119903)
A5
If any soil layers are found within zone 1 (sensitive soils) then caution should be exercised as these
structured clays are prone to instability collapses and difficulties in construction and performance
Another indicator of zone 1 sensitive soils is when the normalized porewater parameter Bq gt 08 When
these criteria are evident the geoengineer should consult with senior geotechnical personnel or the chief
engineer for guidance
Figure A7 Delineation of soil behavioral type zones using CPT material index Ic
A-13
Stiff overconsolidated clayey sands and sandy clays soils of zone 8 (15 lt Fr lt 45) and zone 9 (Fr gt
45) can be identified from the following criterion
Zones 8 and 9 0020)1(00030)1(0050
12
rr
tnFF
Q A6
The remaining soil types are identified by the CPT material index Zone 2 (organic clayey soils Ic ge 360)
Zone 3 (clays to silty clays 295 le Ic lt 360) Zone 4 (silt mixtures 260 le Ic lt 295) Zone 5 (sand mixtures
205 le Ic lt 260) Zone 6 (clean sands 131 le Ic lt 205) and Zone 7 (gravelly to dense sands Ic le 131)
The red dashed line at Ic = 260 represents an approximate boundary separating drained (Ic lt 260) from
undrained behavior (Ic gt 260)
Once the specific zone has been assigned to a layer a visual representation can be made to show
either the zone number or colorization so that the predominant layers and soil types can be
realized
A123 Probabilistic Soil Types from CPT
The inference of soil type has been addressed using probabilistic methods as discussed by Abu-
Farshakh et al (2008) and Tumay et al (2013) Here the CPT readings can be post-processed to
provide the percentage of sand silt and clay components with depth The output is consistent
and compatible with the current MnDOT soil classification chart (Figure A8) which is similar in
content to that used by the USDA
A-14
Figure A8 Chart for soil classification system by MnDOT
A13 Identifying Soft Organic Soils
The identification of soil type using cone penetration tests (CPT) is usually done on the basis of empirical
soil behavioral type (SBT) charts that use the measured readings (qt fs andor u2) While there at least
25 different sets of charts available (ie Hegazy 1998 Eslami et al 2017 Valsson 2016) one of the most
widely used and popular is the 9-zone SBT system that uses three normalized cone readings Qtn Fr and
Bq (Robertson amp Cabal 2007 Lunne et al 1997) The system is denoted as SBTn and plotted on charts in
terms of Q vs F and Q vs Bq as shown in Figure A9
A-15
1
10
100
1000
01 1 10
Norm
aliz
ed
Con
e T
ip R
esis
tan
ce
Q
Normalized Friction Fr = 100 fst(qt - svo) ()
9-Zone Soil Behavioral Type Chart
Gravelly Sands(zone 7)
Sands
(zone 6)
Sandy Mixtures(zone 5)
Silt Mix(zone 4)
Clays
(zone 3)
Organic Soils(zone 2)
Sensitive Clays
and Silts(zone 1)
Very stiff OC clayto silt (zone 9)
Very stiffOC sandto clayey
sand(zone 8)
1
10
100
1000
-06 -04 -02 00 02 04 06 08 10 12 14
No
rmal
ized
Co
ne
Tip
Res
ista
nce
Q
Normalized Porewater Bq = Du2(qt-svo)
Clays(zone 3)
Sensitive(zone 1)Organic
(zone 2)
Gravelly Sands(zone 7)
Sands(zone 6)
Figure A9 CPT charts for SBTn system (a) Q versus F (b) Q versus Bq
Of particular concern in geotechnical site characterization is the proper identification of soft
organic soils such as organic clays and silts (OL and OH) and peats (Pt) These soils often cause
difficulties and problems during construction and performance of civil engineering works because
of bio-degradation long-term creep excessive settlements andor other issues
The SBTn system has a category labelled Zone 2 recognizes organic soils However several recent
studies have noted that data from CPTu soundings do not always fall within the bounds
delineated using Zone 2 even though laboratory tests and field inspections clearly note the
presence of organic soils Lab tests to classify organic soils includes loss on ignition (LOI gt 5)
and two pairs of Atterberg limits testings per ASTM standards as well as the visual-manual
identification of strong organic odor and smell and dark color notably black and dark gray
Results by Mlynarek et al (2014) show that organic soils in Poland are not adequately identified by the
Q-F chart as seen in Figure A10 Moreover studies by Zawrzykraj et al (2017) found that neither the Q-
A-16
F and Q-Bq plots worked well to recognize organic soils and peats in Poland Nejaim et al (2016) found
that Brazilian soft organic clays were misclassified as Zone 3 (clays) rather than Zone 2 (organic clays)
Similarly a recent study on CPTu data from organic clays located at 24 sites in the USA Sweden Mexico
Brazil and Australia have been compiled (Agaiby 2018) These data are shown to generally avoid falling
with the Zone 2 boundaries in either Q-F or Q-Bq charts thus are not recognized during these soundings
as indicated by Figure A11 The exception in this case is the Mexico City clay which is correctly identified
however the other 23 sites are generally not properly recognized and categorized Additional findings by
Missiaen et al (2015) found that the CPTu results in Belgian soils also did not identify organic clays and
peats in the non-normalized version of these charts that uses Qt versus friction ratio (Rf = 100middotfsqt) as
detailed by Robertson amp Cabal (2007)
Figure A10 Soil behavioral chart with superimposed CPTu data from Polish organic soils
(from Mlynarek et al 2014)
A-17
Figure A11 Paired SBTn charts with CPTu data from 24 organic clay sites indicating non-compliance
with Zone 2 (organic soils)
(compiled by Agaiby 2018)
A14 Simplified Yield Stress Evaluation in Clays by CPTu
From the development of a hybrid formulation of spherical cavity expansion (SCE) theory with critical-
state soil mechanics (CSSM) the effective preconsolidation stress or yield stress (σp) of clays can be
expressed in terms of three piezocone parameters (a) net cone tip resistance qnet = qt - σvo (b)
measured excess porewater pressure Δu2 = u2 - u0 and (c) effective cone resistance qE = qt - u2 Details
on the SCE-CSSM solutions are given elsewhere (Mayne 1991 Mayne 2017 Agaiby amp Mayne 2018)
A-18
For normal and well-behaved clays which include inorganic fine-grained soils such as clays and silts
of low sensitivity a set of simple relationships can be derived By adopting characteristic default values
for effective friction angle ϕ = 30deg and rigidity index (IR = 100) linear expressions for the effective
preconsolidation stress are obtained
prime 120590119901 = A7 033 ∙ 119902119899119890119905
prime 120590119901 = 053 ∙ ∆1199062 A8
prime 120590119901 = 060 ∙ 119902119864 A9
A141 Application to soft Chicago clay
An example of a common geomaterial in this category includes the infamous soft Chicago clays
that were deposited in a glacio-lacustrine environment (Cho amp Finno 2010) The soft clay has a
mean water content of 20 liquid limit of 38 and plasticity index PI= 12 Results of CPTu tests
conducted at the national geotechnical experimentation site (NGES) at Northwestern University
are shown in Figure A12 (Ouyang amp Mayne 2017) As seen in the profile a thin soft clay resides
between depths of 9 to 105 m with a thicker soft clay layer in the range of 12 to 18 m
A-19
0
2
4
6
8
10
12
14
16
18
20
00 02 04 06 08 10 12D
ep
th (
m)
CPTu qt and u2 (kPa)
sand (SP)
Blodgett
Park Ridge
Deerfieldsoft clay
ClayLayerof Study
Northwestern UnivNational Test Site
Figure A12 Results of CPTu sounding at the NGES at Northwestern University Illinois
Post-processing of the CPTu data using Equations A7 A8 and A9 show good agreement in evaluating the
profile of effective yield stress in this clay layer compared with results from one-dimensional
consolidation tests made on undisturbed samples taken at the site
A-20
10
11
12
13
14
15
16
17
18
19
20
0 100 200 300 400 500 600 700 800 900 1000
De
pth
(m
)Effective Yield Stress sp (kPa)
Consols
033 qnet
053 Du2
060 qE
svo
Figure A13 Evaluation of effective yield stresses at NWU using consolidation tests and CPTu
A142 Application to soft Bay Mud California
Another example of a well-behaved fine-grained soil is the San Francisco Bay Mud which is a soft clay
deposit in northern California Results of a representative CPTu are shown in Figure A14 and indicate the
soil profile at a site in eastern San Francisco (Hunt et al 2002) For this site index values include 70 lt
wn lt 90 60 lt LL lt 80 and 35 lt PI lt 45 Post-processing the CPTu data using Equations A7 A8 and
A9 show good agreement with each other as well as with the results of consolidation tests as seen in
A-21
Figure A15 The consolidation tests were performed using a CRS device as reported by Pestana et al
(2002)
San Francisco Bay Mud (Pestana et al 2002)
0
5
10
15
20
25
30
35
40
0 1000 2000 3000
De
pth
(m
ete
rs)
Cone Resistance qt (kPa)
0
5
10
15
20
25
30
35
40
0 10 20 30 40 50 60
Sleeve Friction fs (kPa)
0
5
10
15
20
25
30
35
40
0 500 1000 1500
Porewater u2 (kPa)
Miscellaneous Fill
Young Bay Mud
Young Bay Mud
Clayey Sand
Miscellaneous Fill
Young Bay Mud
Young Bay Mud
Figure A14 Representative CPTu in San Francisco Bay Mud (data from Hunt et al 2002)
A-22
San Francisco Bay Mud (Pestana et al 2002)
0
5
10
15
20
25
30
0 100 200 300 400 500 600
De
pth
(m
ete
rs)
Preconsolidation Stress sp (kPa)
svo
033 qnet
053 Du2
060 qE
CRS
svo
Figure A15 Evaluation of effective yield stresses in Bay Mud from consolidation tests and CPTu
A15 CPTu Screening of Soft Organic Clays
For soft organic clays the Equations A7 A8 and A8 will not apply thus can serve as a means to
screen such geomaterials from the soil profile Two examples of CPTu soundings in soft organic
clays are presented to illustrate the approach Additional case records and documentation of
CPTu data in organic clays are given in Agaiby (2018)
A151 Soft organic alluvial soils in Washington DC
Results of in-situ tests including CPTu soundings in soft organic clayey silts along the Potomac River at the
Anacostia Naval Air Station are presented by Mayne (1987) Here the soft soils are categorized as organic
clayey silts (OH) per the Unified Soils Classification Systems (USCS) with mean values (and plus and minus
A-23
one standard deviations) of wn = 68 plusmn 16 LL = 83 plusmn 25 and PI = 37 plusmn 17 A representative
piezocone sounding is depicted in Figure A16 For the post-processing of CPTu data the use of the Q-F
and Q-Bq graphs do not find any points within the Zone 2 category for organic soils When the data are
evaluated to determine the profile of effective yield stress the Equations A7 A8 and A9 do not agree as
shown by Figure A17
0
5
10
15
20
25
0 2000 4000 6000
Dep
th (
met
ers)
Cone Resistance qt (kPa)
0
5
10
15
20
25
0 20 40 60 80 100
Sleeve Friction fs (kPa)
0
5
10
15
20
25
0 200 400 600
Pore Pressure u2 (kPa)
Figure A16 Representative CPTu in soft organic clay along the Potomac River DC
A-24
0
5
10
15
20
25
0 3 6 9D
epth
(m
eter
s)Soil Behaviour Type Zone
Q and Fchart
0
5
10
15
20
25
0 100 200 300 400 500 600
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
033qnet
053 Du2
06 qE
Oed
Figure A17 Post-processing of CPTu data from Anacostia-Bolling (a) SBTn zones using CPT material
index Ic (b) CPTu screening equations for yield stress
Note that the hierarchy of the results from Equations A7 A8 and A9 show that the CPTu
evaluations for preconsolidation stress in soft organic clays indicates
[A10] σp = 053 Δu2 lt 033 qnet lt 060 qE
A152 Soft organic peaty clay in Saint Paul MN
Results of CPTu tests were collected during a training exercise underneath I-35E just northeast
of Saint Paul MN in 2007 All three MnDOT CPT rigs available at the time were used to obtain
in-situ test data at the site The site is well known as having soft organic soils and samples were
obtained from three borings made at the site (Boring ID T523 T542 and T518) Boring logs
A-25
indicate the presence of highly organic silt loams with marl and spongy organic clayey silts
From 12 samples the natural water contents ranged from 30 to 191 with a mean of 112 plusmn
58
A representative piezocone sounding from the site (MnDOT CPTu ID F22Y0703C) is used for
this illustration and is presented as Figure A18 Post-processing these data using the SBTn
algorithms and CPT material index show that the soils classify primarily as Zone 3 (clays) and
Zone 4 (silty mix) thus do not identify the geomaterials properly as Zone 2 (organic soils)
0
5
10
15
0 500 1000 1500 2000
Dep
th (
met
ers)
Cone Tip Resistance
qt (kPa)
0
5
10
15
0 10 20 30 40 50 60
Sleeve Friction
fs (kPa)
0
5
10
15
-100 0 100 200 300 400
Porewater Pressure
u2 (kPa)
Figure A18 MnDOT CPTu sounding at the I-35E test site near St Paul Minnesota
A-26
1
10
100
1000
01 1 10
Norm
aliz
ed
Con
e T
ip R
esis
tan
ce
Q
Normalized Friction Fr = 100 fst(qt - svo) ()
9-Zone Soil Behavioral Type Chart
St PaulGravelly Sands(zone 7)
Sands
(zone 6)
Sandy Mixtures(zone 5)
Silt Mix(zone 4)
Clays
(zone 3)
Organic Soils(zone 2)
Sensitive Clays
and Silts(zone 1)
Very stiff OC clayto silt (zone 9)
Very stiffOC sandto clayey
sand(zone 8)
1
10
100
1000
-06 -04 -02 00 02 04 06 08 10 12 14
No
rmal
ized
Co
ne
Tip
Res
ista
nce
Q
Normalized Porewater Bq = Du2(qt-svo)
Clays(zone 3)
Sensitive(zone 1)Organic
(zone 2)
Gravelly Sands(zone 7)
Sands(zone 6)
Figure A19 Post-processing St Paul sounding using the 9-zone SBTn classification system
Using the recommended approach Figure A20 shows the CPT-evaluated yield stress profiles from
equations A7 A8 and A9 Here again the three estimates of σp do not agree but show the same
hierarchy as detailed in Equation A10
A-27
0
5
10
15
20
0 50 100 150 200 250 300 350 400
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
06 qeff svo
033qnet Consols
054 Du2
Figure A20 CPTu screening of soft organic soils at St Paul test site
A16 CPTu Evaluations of Yield Stress in Soft Organic Soils
Once the presence of soft organic clays and silts is identified using the aforementioned CPT
screening algorithms the evaluation of effective yield stress is conducted using the procedure
outlined in Chapter 2 of this MnDOT manual For soft organic soils an exponent of m = 090 is
recommended (Mayne et al 2009) Therefore the preconsolidation stress is obtained from
(units of kPa)
prime = A11 120590119901 033 ∙ (119902119899119890119905)090
The algorithm can be expressed in dimensionless form by
prime = A12 120590119901 033 ∙ (119902119905 minus 120590119907119900)119898prime ∙ (120590119886119905119898100)1minus119898prime
A-28
where σatm = reference stress (= 100 kPa = 147 psi) and m = 090 for soft organic silts and clays
The approach is applied to the two former example case studies in Figure A21 (Washington DC)
and Figure A22 (St Paul MN) respectively showing good agreement with benchmark results
taken from laboratory consolidation tests on undisturbed samples of these sites in both cases
0
5
10
15
20
25
0 100 200 300 400 500 600
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
033qnet^090
Oed
Figure A21 Profiles of yield stress from consolidation tests and CPTu method for organic clayey silts at
Anacostia-Bolling site in Washington DC
A-29
0
5
10
15
20
0 20 40 60 80 100 120 140 160
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
CPT
Consols
120590119901prime = 033 119902119899119890119905
090 in kPa
Figure A22 Profiles of yield stress from consolidation tests and CPTu method for organic clayey silts at
MnDOT test site near Saint Paul Minnesota
A17 Soil Unit Weight
As shown by Figure A23 total soil unit weight (t) can be estimated from the CPT sleeve friction resistance
(Mayne 2014) The equation can be expressed by
120574119905 = 120574119908 ∙ [122 + 015 ∙ 119897119899 (100 lowast 119891119904120590119886119905119898 + 001)] A13
where w = unit weight of water and satm = atmospheric pressure
A-30
Figure A23 Evaluation of soil unit weight from CPT sleeve friction
A18 Effective Stress Friction Angle
Sands
The effective stress friction angle () is one of the most important soil properties as it governs the strength
of geomaterials as well as affects soil-pile interface and pile side friction While an effective cohesion
intercept (c) can also be considered this is usually reserved for cemented or bonded geomaterials or
unsaturated soils and may lose its magnitude with time ageing or with prolonged environmental
changes
For clean quartz to silica sands where porewater pressures are essentially hydrostatic (Bq = 0) the
following expression has been calibrated with triaxial compression test results from undisturbed sand
samples and normalized cone resistances as presented in Figure A24 (Mayne 2007 2014)
A-31
120601prime(119889119890119892) = 176deg + 110deg 119897119900119892 (1199021199051) A14
Figure A24 Evaluation of effective stress friction angle in quartz-silica sands from CPT
Note Relationship applies to drained soil behavior when Ic lt 26 andor Bq lt 01
where qt1 is an earlier form of stress normalized cone tip resistance for sands given by
(119902119905 120590119886119905119898) 1199021199051 = A151 05 prime (120590119907119900120590119886119905119898)
A-32
In terms of units of bars this is expressed simply as (in units of bars)
119902119905 1199021199051 = 05 A152 prime ) (120590119907119900
Recently Robertson amp Cabal (2015) recommended the use of the modified form of normalized cone
resistance (Qtn) in Equation A10
120601prime (119889119890119892) = 176deg + 110deg 119897119900119892 (119876119905119899) A16
In sands qnet asymp qt since the overburden term is small relative to the cone tip resistance Figure A25 shows
that the two normalizations give very comparable results
Figure A25 Comparable values from qt1 and Qtn normalization schemes for CPTs in sands
Clays and Silts
A-33
In the case of soft to firm intact clays and silty clays the effective friction angles are determined from the
normalized cone resistance and porewater pressure parameters (Senneset et al 1989 Mayne 2016) as
shown in Figure A26 The exact solution when the angle of plastification = 0 is given as
qBQ
)tan1(tan61
1)tanexp()245(tan 2
A17
Approximate NTH Solution for from CPTu
qBQ
)tan1(tan61
1)tanexp()245(tan2
Approx asymp 295deg∙Bq0121∙[0256 + 0336∙Bq + log Q ]
1
10
100
20 25 30 35 40 45
Co
ne
Res
ista
nce
Nu
mb
er Q
(degrees)
Bq
010204060810
c = 0 = 0deg
Theory (dots)
Approximation (lines)
Theory
Figure A26 Evaluation of effective stress friction angle in clays and silts from CPTu results
Note Relationship applies to undrained soil behavior when Ic ge 26 andor Bq ge 01
which can be approximately inverted into the form (Mayne 2007)
A-34
0121 120601prime = 295deg ∙ 119861119902 ∙ [ 0256 + 0336 ∙ 119861119902 + log(119876)] A18
This algorithm is specifically applicable for the following ranges of porewater pressure parameter (01 le
Bq lt 1) and effective stress friction angles (20deg le lt 45deg)
A19 Stress History
The stress history can be characterized by an apparent yield stress or preconsolidation stress (sp) as well
as by its normalized and dimensionless form YSR = spsvo = yield stress ratio The YSR is in effect the
same as overconsolidation ratio (OCR) however now generalized to accommodate mechanisms of
preconsolidation that occur beyond just erosion glaciation and removal of overburden stresses but also
due to ageing desiccation repeated cycles of wetting-drying bonding repeated freeze-thaw cycles
groundwater changes and other factors
A-35
A generalized approach for evaluating the yield stress or preconsolidation in soils using net cone
resistance has been formulated as presented in Figure A27 (Mayne 2015) Yield stress in soils from CPT
10
100
1000
10000
10 100 1000 10000 100000
Yie
ld S
tre
ss
s
y
(kP
a)
Net Cone Resistance qt - svo (kPa)
Blessington
General Trend
sy = 033(qt-svo)m
Fissured Clays m = 11
Intact clays m = 10 Sensitive clays m = 09
Silt Mixtures m = 085Silty Sands m = 080
Clean Sands m = 072
m = 11 10 09
085
080
072
Units kPa
Fissured Clays
Figure A27 Evaluation of yield stress or preconsolidation stress in soils from CPT
The algorithm can be expressed in dimensionless form by
1minus119898prime prime 120590119886119905119898 120590119901 = 033 ∙ (119902119905 minus 120590119907119900)119898prime
( ) A19 100
where m = exponent depends on soil type m = 1 (intact clays) 085 (silts) 080 (sandy silts to silty sands)
and 072 (sands) For fissured clays the exponent m may be 11 or higher depending upon the age of the
A-36
formation and degree of jointing and discontinuities Note that fissured clays can be identified when Ic lt
26 and Bq lt 01
The exponent m has also been calibrated with CPT material index as presented in Figure A28
An algorithm to express this relationship for non-fissured soils is given by
25)652(1
2801
cIm
A20
This allows the post-processing of CPT to automatically choose the appropriate exponent m for
a layer by layer analysis
Figure A28 Yield stress exponent m in terms of CPT material index Ic
A-37
A110 Lateral Stress Coefficient
The horizontal geostatic state of stress is represented by the lateral stress coefficient K0 = shosvo
commonly referred to as the at-rest condition The magnitude of K0 for soils that have been loaded and
unloaded can be approximately estimated from
119870119900 = (1 minus 119904119894119899120601prime) ∙ 119874119862119877119904119894119899120601prime A21
Data compiled from in-situ K0 measurements using self-boring pressuremeter tests (SBPMT) total stress
cells (TSC) or push-in spade cells andor laboratory methods (instrumented consolidometers triaxials
andor suction measurements) on a variety of clays silts and sands have been compiled and reported by
Ku amp Mayne (2015) as shown in Figure A29 verifying Equation A15 as a means for evaluating K0 in soils
Figure A29 Relationship between lateral stress coefficient K0 and YSR or OCR in soils (Ku amp Mayne
2015)
A-38
A111 Undrained Shear Strength
The loading of soils can occur under conditions of being fully drained (Du = 0) partially drained or full
undrained (DVV0) where Du = excess porewater pressures (above hydrostatic) and DVV0 = volumetric
strain Further details on specific stress paths are best explained in terms of critical state soil mechanics
or CSSM (Mayne et al 2009 Holtz Kovacs amp Sheahan 2011) The prevailing drainage conditions depend
upon the rate of loading and permeability characteristics of the soil Normally in sands that are pervious
and exhibit high permeability a drained response occurs Exception to this may occur in loose sands
during fast earthquake loading resulting in soil liquefaction In clays that exhibit low permeability a fast
rate of loading will result in undrained loading at constant volume This in fact is a temporary and
transient condition often termed short term loading In the long term eventually porewater pressures
will dissipate (albeit slowly) and a drained response will prevail thus termed long term loading
The overall soil strength is controlled by the effective stress strength envelope Most commonly this is
represented by a simple linear relationship termed the Mohr-Coulomb criterion where the maximum
shear stress (max called the shear strength) is given as
= 119888prime + 120590prime ∙ 119905119886119899 120601prime A22 120591119898119886119909
where c = effective cohesion intercept s = effective normal stress and = effective stress friction angle
As a starting point values of c = 0 and = 30deg can be adopted for all soil types at least until the specific
soil type geologic formation and results from high-quality laboratory or field test data are available
Peak Undrained Shear Strength from Stress History
Using simplified critical state soil mechanics the peak undrained shear strength (max = su) can be
evaluated from the effective stress strength envelope (c = 0 effective friction angle ) and stress history
(ie OCR) in the form
prime DSS 119904119906 = (119904119894119899120601prime2) ∙ (119874119862119877)Λ ∙ 120590119907119900 A23
A-39
which corresponds to a simple shear mode Figure A30 presents the aforementioned relationship
together with data from 17 different clays tested in simple shear
Figure A30 Normalized undrained shear strength from simple shear tests on various clays
versus YSR or OCR
For the triaxial compression (TC) mode the equation would give slightly higher strengths that are
calculated from
prime TC 119904119906 = (1198721198882) ∙ (1198741198621198772)Λ ∙ 120590119907119900 A24
where Mc = (6 middot sin)(3-sin)
Peak Undrained Shear Strength from CPTu
A more direct approach to assessing undrained shear strength is via bearing capacity theory
whereby
A-40
A25 kt
votu
N
qs
s
where Nkt = bearing factor that depends on mode of shearing sensitivity of the clay degree of
overconsolidation and other factors For soft offshore clays the back figured Nkt from data collected at
14 well-documented sites (Low et al 2010) determined mean values based on mode of shearing Nkt =
119 (triaxial compression) Nkt = 136 (simple shear) and Nkt = 133 (vane)
A recent study of 51 clays that were tested by both field piezocone and laboratory CAUC triaxial tests
showed that essentially Nkt = 12 for intact clays and clayey silts of low to medium sensitivity (Mayne
Peuchen amp Baltoukas 2015) Figure A31 shows the slightly different trends for offshore versus onshore
clays For sensitive clays a lower value of Nkt = 10 would be appropriate and for fissured clays a higher
value of Nkt would be in the range of 20 to 30
A-41
A-42
Figure A31 Database from 51 clays with CPT qnet versus lab measured triaxial shear strength
(Mayne Peuchen amp Baltoukas 2015)
For soft to firm clays an alternate means to evaluate undrained shear strength is via the excess porewater
pressures ( u = u2 - u0) from the following
u
uN
us
D
D 2
A26
where N u = porewater bearing factor also dependent on the aforementioned factors The study by
Low et al (2010) found N u = 59 (triaxial compression) N u = 69 (simple shear) and N u = 71 (vane
shear)
A third alternate is found from the effective cone resistance (qE = qt - u2) whereby
kE
tu
N
uqs 2
A27
where NkE is a bearing term for this approach Mayne et al (2015) indicated a mean value of NkE = 8 for
soft-firm intact clays
Remolded Undrained Shear Strength from CPT
The remolded undrained shear strength (sur) is obtained from either field vane tests lab mini-vane shear
tests or lab fall cone devices This affords the evaluation of the clay sensitivity (St) which is defined as the
ratio of peak to remolded strengths at a given water content
119904119906(119901119890119886119896) 119878119905 = frasl A28 119904119906119903
For the CPT the sleeve friction has been noted to give values that are comparable to the
remolded undrained shear strength (Powell amp Lunne 2015) Therefore
asymp 119891119904 A29 119904119906119903
A112 Ground Stiffness and Soil Moduli
The stiffness of the ground can be represented by several geoparameters including the compressibility
indices (Cr Cc Cs) spring constants (kv) and moduli Regarding the latter there are several moduli that
are used in geotechnical engineering including shear modulus (G and Gu) Youngs modulus (E and Eu)
constrained modulus (D) bulk modulus (K) resilient modulus (MR) and subgrade reaction modulus (ks)
A-43
Soil Modulus
The definition of modulus is taken as E = DsD with Figure A32 showing a typical deviator stress (q = s1
- s3) versus axial strain (a) curve Theoretical interrelationships between the elastic moduli G E D and
K have a dependence on the Poissons ratio v The value of Poissons ratio can be taken as vu = 05 for
undrained loading (ie constant volume) while for drained loading which is accompanied by volumetric
strains a value of v = 02 may be used If we adopt the reference modulus as E then the
interrelationships with the other elastic moduli are given by
a Reference Stiffness 119864prime = drained Youngs modulus A30
119864prime
b Shear Modulus 119866prime = A31 [2(1+120584prime)]
119864prime (1minus120584prime) c Constrained Modulus 119863prime = A32
[(1+120584prime)(1minus2120584prime)]
119864prime
d Bulk Modulus 119870prime = A33 [3middot(1minus2120584prime)]
A-44
Figure A32 Representative stress-strain-strength curve for soils in triaxial compression
For the extreme case when = 0 in fact D = E When = 02 the two moduli are only 10
different D = 11middotE thus the constrained modulus and drained Youngs modulus are often
considered somewhat interchangeable
The resilient modulus (MR) applies to pavement analysis and design most commonly measured by cyclic
triaxial testing under repeated load applications In fact MR is a special case of Youngs modulus that
relates to the small strain stiffness measured in the nondestructive range but has developed permanent
plastic strains after many cycles of loading (Brown 1996 Dehler amp Labuz 2007)
The subgrade reaction modulus is actually a combined soil-structural parameter as its value
depends on the ground stiffness and the size of the loaded element The subgrade modulus is
defined as ks = q where q = applied stress and = measured deflection In terms of elasticity
solutions the deflection of a flexible circle of diameter d is given by = qmiddotdmiddot(1-2)E Therefore
119864prime
119896119904 = A34 [119889middot(1minus1205842)]
which has units of kNm3 or pcf
Table A2 summarizes several selected studies towards the approximate evaluation of these moduli
obtained directly from measured CPT data Note however that the results from in-situ penetrometer
tests generally represent a peak strength as indicated in Figure A32 Thus the measured cone tip
resistance (qt) reflects the top of the stress-strain curve either the undrained strength in clays (su) or the
effective friction angle () of sands or even a strength intermediate between these two values Thus
A-45
Source Soils Studied Reference Modulus
Findings
Kulhawy amp Mayne 1990
8 stiff OC clays Lab constrained modulus
D asymp 825 middot (qt - svo)
Mohammed et al (2000)
Resilient modulus
Abu-Farsakh 2004
7 Louisiana sites
Lab constrained modulus
D asymp 358 middot (qt - svo)
Mayne (2007b)
All soil types Constrained moduli from lab consolidation
First-order estimate D asymp middot(qt - svo)
Robertson (2009)
All soil types Constrained modulus from consolidation
D = m middot (qt - svo)
when Ic gt 22
1 m = Qtn when Qtn le 14
2 m = 14 when Qtn gt 14
when Ic lt 22 then
= 003 middot 10055 Ic + 168 m
Liu et al (2016)
16 clay sites in China
Resilient modulus MR = (146qt 053 + 1355 fs
14 + 236)244
(all in MPa)
Casey et al (2016)
8 clays Triaxial tests for undrained Youngs modulus
Eu(svc)07 = 316 - 23 middot LL
where Eu and svc (MPa)
and LL = liquid limit ()
qt is probably not the best means to obtain the slope of the stress-strain curve Instead the SCPTu that
provides the initial tangent shear modulus (Gmax) from the shear wave velocity (Vs) is a better choice
Table A2 Selected Modulus Relationships with Cone Penetration Test Measurements
A-46
Nonlinear Modulus
The small-strain shear modulus (Gmax or G0) represents the initial stiffness of all soils and rocks It is the
beginning of all stress-strain-strength curves for geomaterials and obtained from elasticity theory
119866119898119886119909 = 1198660 = 120588119905 middot 1198811199042 A35
where Vs = shear wave velocity t = tga = total soil mass density t = total soil unit weight and ga =
gravitational acceleration constant (98 ms2)
From a general viewpoint on stiffness the shear modulus G of soil is defined as the slope of shear stress
versus shear strain G = DDs for a tangent definition and by G = s as a secant definition The shear
modulus is related to its associated Youngs modulus E = 2G(1+v) Both moduli are in fact highly
nonlinear ranging from a maximum value at the small-strain stiffness (Gmax = G0 = t middotVs2 where t =
total soil mass density and Vs = shear wave velocity) to intermediate G values at medium strains (asymp 1)
to low values at peak strength As such a variety of algorithms and formulae have been developed to
represent either a partial range or the full stress-strain-strength behavior of soils over a range of
interests (eg Mayne 2005) In these formulations a variable number of input parameters may be
required in order to produce a stress-strain-strength curve
A modified hyperbolic form suggested by Fahey amp Carter (1993) has favorable attributes in that the
modulus reduction factor can be established with only a single variable This allows the initial stiffness
to the be small-strain stiffness (G0) that is reduced in terms of level of mobilized strength eg max =
qqmax = 1FS which is simply the reciprocal of the calculated factor of safety (FS) The magnitude of
secant shear modulus (G) corresponding to the particular level of loading is given by
119866 = 119872119877119865 middot 1198660 = (1198661198660) middot 1198660 A36
A-47
where the MRF = modulus reduction factor determined as
1 minus (120591 )119892 119872119877119865 = (1198661198660) = frasl A37 120591119898119886119909
and g = empirical exponent term Figure A33 shows the relationships for normalized shear stress ()
versus shear strain (s) and corresponding normalized secant shear modulus (G = s) with shear strain
over a range of exponent values 01 le g le 10 For a discussion of tangent G please see Fahey amp Carter
(1993)
Using laboratory data from resonant column-torsional shear tests and triaxial specimens with local
strain measurements for Gmax reference values modulus reduction curves for a selection of sands and
clays are presented in Figure A34 The data indicate that the exponent g falls within the range 02 lt g lt
05 for many soils tested drained and undrained A mean value of g = 03 is recommended for
preliminary site investigations and designs until additional information can be obtained
The value of max is the shear strength of the soil generally taken as either (1) drained (c = 0) max =
svo middot tan or (2) undrained where max = su = undrained shear strength as discussed earlier Thus each
shear stress () can be associated with its shear modulus (G) and the relevant shear strain is found from
s = G This allows for generation of nonlinear stress-strain-strength curves at all depths from SCPTu
data in clays sands and mixed soil types
A-48
Modified Hyperbola (Fahey amp Carter 1993 Canadian Geotech J) Coefficient Term f = 1
``
0
01
02
03
04
05
06
07
08
09
1
000001 00001 0001 001 01 1
No
rmal
ized
GG
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
0
01
02
03
04
05
06
07
08
09
1
0 01 02 03 04 05 06 07 08 09 1
No
rmali
zed
GG
ma
x
Mobilized Shear Stress max
g =1
07
05
03
02
01
0
01
02
03
04
05
06
07
08
09
1
00001 0001 001 01 1
No
rmali
zed
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
0
01
02
03
04
05
06
07
08
09
1
0 1 2 3 4 5 6 7 8 9 10
No
rmali
zed
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
Reduction Factor
GGmax = 1- (max)g
Figure A33 Normalized modulus (GGmax) and normalized shear stress (max) versus shear strain (s)
for modified hyperbolic relationship of Fahey amp Carter (1993)
A-49
0
01
02
03
04
05
06
07
08
09
1
0 01 02 03 04 05 06 07 08 09 1
Mo
du
lus
Re
du
cti
on
G
Gm
ax
or
EE
ma
x
Mobilized Strength max or qqmax
NC SLB Sand
OC SLB Sand
Hamaoka Sand (e=0625)
Hamaoka Sand
Toyoura Sand e = 067
Toyoura Sand e = 083
Ham River Sand
Ticino Sand
Kentucky Clayey Sand
Kaolin
Fujinomori Clay
Pietrafitta Clay
London Clay
Vallericca Clay
Pisa Clay
Pisa Clay
Pisa Clay
Pisa Clay
Kiyohoro Silty Clay
Thanet Clay
Exp g = 02
Exp g = 03
Exp g = 04
Exp g = 05
MRF = 1 - (qqmax)g
Figure A34 Modulus Reduction Factors (MRF) with mobilized strength of clays and sands
A113 Coefficient of Consolidation
The rate at which foundation and embankment settlements occur as well as the dissipation of excess
porewater pressures is controlled by the coefficient of consolidation (cv) The magnitude of cv is also
required in the design of vertical wick drains that can be installed in soft ground to expedite the time for
consolidation Using the results of CPTu dissipation tests which measure the rate at which the u2 readings
vary with time the in-situ profile of cv can be evaluated Often the piezocone-interpreted values of cv
are validated by comparison with results from laboratory one-dimensional consolidation tests (eg Abu-
Farsakh MY 2004) Alternatively a better method is to cross-check the values with the measured full-
scale performance of instrumented embankments that are constructed over the ground and the recorded
time-rate of consolidation and settlements can provide the best cv for that site (Abu-Farsakh MY et al
2011)
A-50
Monotonic Dissipation Tests
An illustrative dissipation curve from piezocone testing at IDT State Highway 95 at an embankment and
bridge crossing is presented in Figure A35 After the CPTu sounding was halted at a depth of 512 feet
the recorded decay of porewater pressures was observed to be monotonic with time until the dissipations
were ended at 1000 seconds During that time the u2 readings decreased from 115 psi to 47 psi At this
location the depth to the groundwater table is about 16 feet thus the calculated equilibrium water
pressure is 15 psi Commonly a characteristic time for piezo-dissipations is taken at 50 degree of
consolidation although other degrees may be adopted By evaluating the value of u2 at 50 as shown in
Figure A35 the characteristic t50 = 404 s can be obtained
0
20
40
60
80
100
120
1 10 100 1000 10000
Pore
wat
er P
ress
ure
u2
(psi
)
Time (s)
CPTU-02E-1 z = 1560 m
uo (hydrostatic)
u2 (initial) = 115 psi
u2 (final projected) = u0 = 15 psi
u2 at 50 = frac12(115+15) = 65 psi
Idaho DOT State Route 95Dissipation at 512 feet
t50 = 404 seconds
Time to reach50 consolidation
A-51
Figure A35 Piezo-dissipation at Sandpoint Bridge Crossing Idaho for depth z = 512 feet illustrating the
evaluation of characteristic time t50
It is also common to plot normalized excess porewater pressures relative to the measured initial Du2
that is obtained during penetration at the constant rate of 20 mms ie DuDui versus time Figure A36
shows a set of piezo-dissipation records for a bridge and embankment site in southern Louisiana
reported by the LTRC While the porewater pressure axis is arithmetic the time scale is often plotted on
either logarithmic or square root scales In either case the t50 is then found from the measured
dissipation curve when DuDui = 050
A-52
Figure A36 Set of monotonic piezo-dissipations taken at various depths for Courtableau Bridge
Site on Louisana State Highway 103 (Abu-Farsakh et al 2011)
Dilatory Dissipation Curves
In a number of situations a monotonic type dissipation does not occur on the onset but instead a dilatory
type curve is observed Here after stopping the push of the penetrometer the recorded porewater
pressures initially begin to rise and eventually reach a peak value then afterwards follow a phase where
the Du values decrease to hydrostatic (Figure A37) A full solution is available for both cases (Burns amp
Mayne 2002) Dilatory responses are most often associated with overconsolidated clays and silts
although occasionally are observed in fine-grained soils with low OCRs
The selection of the characteristic t50 value may found by using a square root of time plot to find the initial
u2 value by projection of the post-peak dissipatory data back to the ordinate axis as shown by the example
in Figure A38 In this example the initial u2 = 480 kPa and then the same procedures in Figure A35may
apply
A-53
Figure A37 Monotonic and dilatory porewater dissipation responses in soils
(after Burns amp Mayne 1998)
Figure A38 Method for obtaining t50 from dilatory dissipation curves
(Schneider amp Hotstream 2010)
A-54
Interpretation of Dissipation Test Data
There are a number of different solutions available for the interpretation of the coefficient of
consolidation (cv) from the piezocone dissipation curves For monotonic type response the strain path
method (SPM) developed at Oxford University (Teh amp Houlsby 1991) is well-recognized while the Georgia
Tech solution by Burns amp Mayne (2002) is based on spherical cavity expansion and critical state soil
mechanics (SCE-CSSM) and handles both monotonic and dilatory curves For both solutions a simplified
procedure can be recommended that relies on the aforementioned t50 value obtained from the measured
field data as given in Table A3 The units for cv are in cm2s as shown or equivalent
Table A3 Recommended procedures for calculating cv from t50 obtained in dissipation tests
Method Equation for coefficient of
consolidation cv
RemarksNotes Eqn
No
SPM
(Teh amp
Houlsby
1991) 50
2)(2450
t
Iac
Rc
v
t50 = measured time to reach 50
dissipation
IR = Gsu = undrained rigidity index
G = shear modulus
su = undrained shear strength
ac = penetrometer radius (ac = 178
cm for 10-cm2 cone ac = 220 for a
15-cm2 size)
A32
SCE-CSSM
(Burns amp
Mayne 2002) 50
7502 )()(0300
t
Iac Rc
v
A33
A-55
Evaluating rigidity index
Both the SPM and SCE-CSSM solutions require an estimate of the in-situ rigidity index (IR = Gsu) of the
soil If the results of SCPTU are available Krage et al (2014) have derived an expression for IR that
depends upon the small-strain shear modulus net cone tip resistance and effective overburden stress
250750
max50
8111
vonet
Rq
GI
s A38
where consistent units are input for Gmax qnet and svo The value of Gmax is obtained from B29 using the
measured shear wave velocity (Vs) and unit weight evaluated from B7
If the shear wave velocity is not measured it can be estimated from the CPT data A number of
methods have been reviewed for CALTRANS by Wair et al (2012) including one that is applicable
to sands silts and clays of low sensitivity that are inorganic and uncemented
119881 (119898119904) = [101 middot 119897119900119892(119902119905) minus 114]167 middot (100 middot 119891119904119902119905)03 A39 119904
where qt is in units of kPa (Hegazy amp Mayne 1995) An alternative relationship is given by Robertson amp
Cabal (2015)
A114 Hydraulic Conductivity
The hydraulic conductivity is a parameter that expresses the flow characteristics of the soil In
geotechnics it is also called the coefficient of permeability (k) and has units of cms or feetday
Through consolidation theory the hydraulic conductivity relates directly to the coefficient of
consolidation
A-56
A40 D
ck wv
where w = unit weight of water and D = constrained modulus
Therefore one approach to evaluating k can be from the site-specific cv obtained from the
dissipation tests using an estimate of D from one of the relationships given in Table A2
An approach developed for soft normally-consolidated soils that uses t50 directly to assess the magnitude
of k is presented in
Figure A39 (Parez and Fauriel 1988) The mean trendline through the wedges of soil type can be
approximated by
251
50 (sec)251
1)(
tscmkh
A41
A-57
Figure A39 Hydraulic conductivity versus dissipation time for 50 consolidation
(Parez amp Fauriel 1988)
A-58
APPENDIX B
List of Figures
Figure B1 Conventional methods vs direct CPT approach to shallow foundation response B-3
Figure B2 Direct bearing capacity relationship for embedded square and strip footings situated on clean
sands (after Schmertmann 1978) B-6
Figure B3 Direct CPT bearing factor Rk as function of footing width B and embedment depth from finite
element analyses (Tand et al 1995) B-6
Figure B4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio and sand
consistency (Eslaamizaad amp Robertson 1996) B-7
Figure B5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp Salgado (2005) in
terms of footing size relative density and base settlement B-8
Figure B6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and normalized
cone tip resistance (after Eslami amp Gholami 2005) B-8
Figure B7 Direct relationship between ultimate bearing stress in clays and measured cone tip resistance
(Schmertmann 1978) B-10
Figure B8 Direct relationship between ultimate foundation bearing stress in clays and net cone tip
resistance (after Tand et al 1986) B-11
Figure B9 Measured load-displacements for each of the five large footings at TAMU (Briaud amp Gibbens
1994) sand site B-12
Figure B10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU (Briaud amp
Gibbens 1994) footings B-13
Figure B11 Characteristic stress versus square root normalized displacements for TAMU footingsB-15
Figure B12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sandB-16
Figure B13 Summary of sand formation factors with corresponding CPT resistancesB-19
Figure B14 Normalized foundation stresses vs square root of normalized displacement for spread
footings on sand (modified after Mayne et al 2012) B-20
Figure B15 Definitions of capacity from three types of observed foundation load test responses
(modified after Kulhawy 2004) B-21
Figure B16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footingsB-22
B-1
Figure B17 Measured load vs displacement curves for 3 shallow foundations on clay at Baytown Texas
(Stuedlein amp Holtz 2010)B-25
Figure B18 Characteristic stress vs square root of normalized displacement for all three foundations at
Baytown site (Stuedlein amp Holtz 2010) B-26
Figure B19 Normalized foundation stress vs square root of normalized displacements B-27
Figure B20 Applied foundation stress vs CPT-calculated stress for 67 footingsB-28
Figure B21 Applied footing stress vs CPT calculated stress on logarithmic scales B-29
Figure B22 Summary graph and equations of direct CPT method for evaluating the vertical B-30
Figure B23 Foundation soil formation parameter hs versus CPT material index Ic B-32
Figure B24 Bearing capacity ratio qmaxqnet versus CPT material index IcB-33
Figure B25 Influence factor for rectangular foundations from elastic theory solution (Mayne amp
Dasenbrock 2017) B-34
List of Tables
Table B1 Direct CPT methods for bearing capacity of footings on clean sandsB-4
Table B2 Direct CPT methods for bearing capacity of footings on clays B-9
Table B3 Summary of large footings soil conditions and reference sources of database B-17
Table B4 Sources for shallow foundation performace B-34
B-2
B1 DIRECT CPT METHOD FOR SHALLOW FOUNDATIONS
B11 Introduction
The analysis of shallow foundations on soils is classically handled as a two-step and two-part set of
calculations involving (a) bearing capacity commonly reliant on limit plasticity solutions and (b)
settlement or more appropriately termed displacements that are assessed via elastic continuum theory
The utilization of cone penetration tests (CPT) provides the necessary geotechnical data for the site-
specific information on the subsurface conditions at the project site including the geostratigraphy and
evaluation of input parameters This is still a viable approach where the CPT readings are interpreted to
give the soil unit weight (γt) effective friction angle (ϕ) undrained shear strength (su) preconsolidation
stress (σp) and elastic moduli (D and E) for analysis
An alternative approach is the use of CPT results to provide direct assessments of bearing capacity
andor settlements A comparison of these two distinctly different and alternate paths is depicted in
Figure B1
Figure B1 Conventional methods versus direct CPT approach to shallow foundation response
A number of available direct CPT approaches for shallow footings are reviewed within this report
Moreover as the entire load-displacement-capacity response of shallow foundations to loading occurs
as a continuous nonlinear phenomenon a single direct CPT solution is presented for the behavior of
B-3
footings on sands (drained) and clays (undrained) The methods discussed herein are specifically to
address the general case of vertical loading of shallow foundations Additional more complex situations
that consider load eccentricity moments inclined forces sloping ground and other facets may be
handled using well-established procedures that are discussed elsewhere (eg Vesić 1975 Kulhawy et
al 1983)
This report focuses on foundations on granular soils andor soils exhibiting drained behavior since
Federal Highway Administration (FHWA) studies have concluded that less than 1 of shallow
foundations for highway bridges are placed on clay soils (Paikowsky et al 2010) One reason for this is
because soft-firm intact clays require a more complicated process that assesses both a short-term
analysis (undrained) as well as long-term analysis (drained) thus requiring undrained bearing capacity
undrained distortion displacements drained bearing capacity and drained primary consolidation
settlements
B12 Direct CPT Methods for Bearing Capacity of Footings on Sands
Classical bearing capacity solutions are based most often in limit plasticity theory albeit can be
formulated from limit equilibrium cavity expansion and numerical methods Because the limit
plasticity solutions are prevalent and adopt total stress analysis they are usually developed as either
fully drained (Δu = 0) or fully undrained (ΔVV = 0) where Δu equals excess porewater pressure and
ΔVV equals volumetric strain Paikowski et al (2010) provides an extensive review of various solutions
There are a number of direct CPT methods for the evaluation of the bearing capacity of footings and
shallow foundations Table B1 lists a variety of these direct CPT approaches for footings on sands In
the direct approach a one-step process is used to scale the measured cone penetrometer readings (ie
measured cone tip resistance (qc) total cone tip resistance (qt) sleeve friction (fs) andor measured
porewater pressure u2)) to obtain the ultimate bearing stress (qult) of the foundation
Table B1 Direct CPT methods for bearing capacity of footings on clean sands
Method Surface Footing Remarks and Embedment Notes
Meyerhof (1956) qult = qc (B12) middotcw qult = qc (B12)(1+DeB)cw
Note for silty sand reduce by 05
cw = 10 dry or moist sand cw = 05 submerged sand
Meyerhof (1974) qult = qc (B + D)40 with stresses in tsf
Presumably dry sands B = fdn width (feet) and D = depth (feet)
Schmertmann (1978)
NA (applies to embedded footings)
Square qult =055σatm(qcσatm)078
Strip qult =036σatm(qcσatm)078
See Figure A2
Embedment applies De gt 05(1+B) for Blt1m De gt 12 m for B gt1m
B-4
Canadian qult = Rk0middotqc Applied to FS = 3 Geotech Society where Rk0 = 03 where FS = factor of (CFEM 199 2) safety
Tand et al (1995) NA qult = Rkmiddotqc + σvo See Figure A3 where Rk = fctn(De B)
Frank and Magnan (1995)
qult = Rk0middotqc where Rk0 = fctn(sand
consistency) Loose Rk0 = 014
Medium Rk0 = 011 Dense Rk0 = 008
qult = Rk1middotqc + σvo where Rk1 = function (De B L amp
sand consistency) Factor Rk1 = Rk0[1+Rk206+04(BL) (DeB)]
and Rk2 = 035 (loose) 050 (medium) and 085 (dense)
Loose sand qc lt 5 MPa
Medium sand 8 MPa lt qc lt 15MPa
Dense sand qc gt 20 MPa
Eslaamizaad amp Robertson (1996)
qult = KΦmiddotqc See Figure A4
See surface footing equation KΦ = function (BDe shape and density)
Lee amp Salgado (2005)
qbL = βbcmiddotqc(AVG)
where qc averaged over distance B
Not addressed See Figure A5 for factor βbc = fctn(B DR
K0 and sB)
beneath the base
Eslami amp Gholami (2005
2006)
See embedded footing solution
qult = Rk1middotqc where Rk1 = function (ratio DeB
and normalized qcσvo) See Figure A6
Measured qc and qcσvo are geometric means over 2B deep
beneath footing
Robertson amp Cabal (2007)
qult = KΦmiddotqc with KΦ = 016
See surface footing equation
Briaud (2007) qult = KΦmiddotqc with KΦ = 023
Based on full-scale tests at Texas AampM
Mayne amp Illingworth
(2010) qult = 018 qc-Mean
Note qc -Mean is averaged CPT cone resistance over depth of
influence z = 15 B deep
Based on 30 footing load tests on 12
sands
Lehane (2013) qult = 016 qc-AVE Note qc-AVE is averaged CPT cone resistance within z (m) = [B (m)
]07
Based on 47 load tests
Table B1 Continued
Notes B = minimum footing width (or diameter) L = footing length De = embedment depth cw = water
table correction σvo = total overburden stress at bearing elevation and σvo = the effective vertical stress
B-5
Figure B2 Direct bearing capacity relationship for embedded square and strip footings
situated on clean sands (after Schmertmann 1978)
Figure B3 Direct CPT bearing factor Rk as function of footing width B and embedment depth
from finite element analyses (Tand et al 1995)
B-6
Figure B4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio
and sand consistency (Eslaamizaad amp Robertson 1996)
B-7
Figure B5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp
Salgado (2005) in terms of footing size relative density and base settlement
Figure B6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and
normalized cone tip resistance (after Eslami amp Gholami 2005)
B13 Direct CPT Methods for Bearing Capacity of Footings on Clays
For direct CPT evaluations of foundation bearing capacity in clays Table B2 summarizes some of the
well-documented methods that are available Generally these assume that the loading is fast enough
and the permeability of the clay is also sufficiently low such that an undrained (ie constant volume)
condition is maintained This is fine for short term loading of foundations particularly for soft to firm
intact clays However the undrained case is not permanent Given sufficient time excess porewater
pressures in the clay will eventually dissipate and hydrostatic conditions will return to equilibrium
During this dissipation phase primary consolidation will occur that results in drained settlements Also
a drained bearing capacity condition will prevail As the CPT is advanced at a constant rate of 20 mms
the recorded readings normally constitute undrained behavior from a direct measurement viewpoint
Thus direct CPT methods are normally focused at an undrained foundation response Nevertheless the
drained footing case can be addressed using CPT results via use of conventional analysis approach and
effective stress limit plasticity solutions that require piezocone penetrometers with porewater pressure
measurements
B-8
B
D)
L
B4060(3501320kc
Many of the earlier solutions were based on data from mechanical penetrometers andor older electric
cones that measure the uncorrected tip resistance (qc) It is now well-recognized that this measurement
must be updated to the corrected cone resistance (qt) because of porewater pressure effects that act on
the mechanics of the load cell and geometry of the particular penetrometer design The results are
particularly significant in clays silts and mixed soils that are soft to firm to stiff where excess porewater
pressures are recorded
Table B2 Direct CPT methods for bearing capacity of footings on clays
Method Footing Comments NotesRemarks
Meyerhof (1974) qult = αbcmiddotqc
where 025 le αbc le 050
Applicable to saturated insensitive clays under short-term loading
Cone tip resistance
measured by
mechanical CPT
Trofimenkov (1974)
Approximation
qult asymp (qc33)09
(Assuming FS = 3)
Strip footings on clays and sandy clays 06m le B le 15m 1m le zemb le 25m
Mechanical CPT
with qc and qult in
kgcm2
Schmertmann (1978)
qult = function (qc and foundation shape)
Relationship shown in Figure A7 for square and strip footings
Cone tip resistance
measured by
mechanical CPT
Tand et al (1986)
qult = Rkmiddot(qc - σvo) + σvo
Factor Rk obtained from Figure A8 accounts for surface and deep footings Separate Rk factor for intact and jointed clays
= (qc1middotqc2)05 qc
where qc1 is geometric mean from bearing elevation to 05B deeper and qc2 is geometric mean from 05B to 15B beneath foundation base
Mix of data from
mechanical CPT qc
and electric CPT qc
LCPC Method
(Frank amp Magnan 1995)
Footing a surface
qult = kc middotqc + σvo
where kc = 032
Embedment case
Note Methods above generally consider undrained bearing capacity
B-9
Figure B7 Direct relationship between ultimate bearing stress in clays and measured cone tip
resistance (Schmertmann 1978)
B-10
Figure B8 Direct relationship between ultimate foundation bearing stress in clays and net
cone tip resistance (after Tand et al 1986)
B14 Direct CPT Assessments of Settlement and Displacements of Shallow Foundations
Methods for evaluating the magnitude of foundation displacements (s) can be found using
elastic theory subgrade reaction methods spring models and numerical simulations The
classical approach to footing settlement calculations on sands is to utilize elastic theory
solutions (eg Poulos amp Davis 1974) which take on the form
119902∙119861∙119868∙(1minus1205842) 119904 = B1
119864119904
where q = applied footing stress and I = displacement influence factor from elasticity theory
The value of I depends upon foundation geometry footing rigidity embedment depth variation of soil
modulus with depth compressible layer thickness and other factors as discussed by Mayne amp Poulos
(1999) For instance for a flexible circular footing of diameter B resting on an infinitely deep soil
formation having a constant modulus Es with depth the influence factor I is equal to 1 for the center
point The results of in-situ field tests such as pressuremeter tests (PMT) standard penetration tests
(SPT) cone penetration tests (CPT) flat dilatometer tests (DMT) and other methods can be used to
ascertain the input value of soil modulus Es
Alternatively direct CPT methods have been investigated to provide a one-step assessment of
foundation displacements Selected methods for direct CPT evaluation of shallow foundation
settlements on granular soils is include DeBeers amp Martens (1957) Meyerhof (1965) DeBeer (1965)
Thomas (1968) Schmertmann (1970) Meyerhof (1974) Berardi amp Jamiolkowski amp Lancellotta (1991)
Robertson (1991) Lutenegger amp DeGroot (1995) Lehane amp Dougherty amp Schneider (2008) Gavin amp
Adekunte amp OrsquoKelly (2009) Mayne amp Illingworth (2010) and Uzielli amp Mayne (2012)
B141 Large Scale Footing Load Tests
The measured load-displacement response of shallow foundations is conducted in the field using either
stepped loading procedures or continuous rate of displacement methods Results from the well-
documented test program at the Texas AampM University (TAMU) site (Briaud amp Gibbens 1999 Briaud
2007) for five spread footings on sand are shown in Figure B9 Each footing clearly shows a nonlinear
B-11
behavior to loading over the range of testing This site is one of the national geotechnical
experimentation sites (NGES) funded by the National Science Foundation (NSF) FHWA and American
Society of Civil Engineering (ASCE)
Figure B9 Measured load-displacements for each of the five large footings at TAMU (Briaud
amp Gibbens 1994) sand site
B142 Generalized Direct CPT Method for Footing Response on Soils
The use of applied foundation stress versus normalized displacement curves (q vs sB) is generalized to
footings on sands silts intact clays and fissured clays A square root plotting of the normalized
displacements (sB) permits a single parameter characterization for each specific soil type that in turn
correlates with the net cone tip resistance The generalization is based on a statistical review of data
from large foundations (B gt 05m) involving 70 full-scale load tests with available CPT data For fine-
grained soils the data are from electronic piezocone tests where the proper correction from qc to qt has
been obtained to allow the best possible results The methodology permits an evaluation of the load-
displacement-capacity response of shallow square footings based on CPTs performed in the four types
B-12
of ground conditions For the TAMU footings the characteristic stress-normalized displacement
response is shown in Figure B10 where all five foundations can be represented by a single curve
Figure B10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU
(Briaud amp Gibbens 1994) footings
B143 Characteristic Stress-Displacement Curves
Fellenius amp Altaee (1994) recommended the concept of a characteristic stress (q) vs normalized
displacement (sB) curve for foundation response on a given soil formation later supported by Decourt
(1999) Briaud amp Gibbens (1999) Lutenegger amp Adams (2003) and Briaud (2007) In this approach the
measured load (Q) vs displacement from individual footings of various sizes that rest on the same soil
conditions collapse to a unified relationship given in Equation B2
119939119943 119956 119954119938119953119953119949119946119942119941 = 119938119943 (119913
) B2
B-13
where af and bf are empirical fitting coefficients (Decourt 1999 Uzielli amp Mayne 2011 2012)
In a review of measured load tests data from large spread footings situated on different sands it has
been suggested that Equation B2 can be reduced to a single parameter expression by use of square root
plotting where the exponent bf is equal to 05 (Mayne amp Illingworth 2010 Mayne et al 2012)
119956 119954119938119953119953119949119946119942119941 = 119955119956radic
119913 B3
where rs = fitted soil parameter from best fit line regression In the case of sands the coefficient rs
represents the supporting soil conditions including particle size relative density and fines content as
well as geologic origin aging and overconsolidation effects
The results from the five TAMU footings are presented in this format in Figure B11 that shows the
formation factor rs = 486 MPa for this sand deposit This single coefficient can be used to express the
nonlinear load versus settlement of any size footing on this sand site Moreover one criterion from
Eurocode for bearing capacity that is used by the European community identifies that stress (or load)
which corresponds to (sB) = 10 That is when the displacement equals ten percent of the footing
width then the capacity has been reached Therefore adopting this criterion for footings on sands
the ultimate bearing stress (qult) for footings on sand can be taken when (sB) equals 010 or when
square root of (sB) equals 0316 In that case this gives qult equal to 0316 middot rs For the TAMU site this
gives qult equal to 0316 times (486) which equals 153 MPa as illustrated by Figure B12
B-14
Figure B11 Characteristic stress versus square root normalized displacements for TAMU
footings
B-15
Figure B12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sand
B15 Footing Database
A database of spread footings and large plates was assembled which considers only full-size shallow
foundations (05 le B le 6 m) that rest on sands silts and clays Sites for the foundations were also subjected to CPT The inclusion solely of large footings are significant because results from small-scale
model footings exhibit scale effects In fact experimental centrifuge work and numerical simulation
studies have shown that bearing capacity factors are size-dependent especially for factor Nγ decreasing
with footing width B (Kimura et al 1985 Cerato amp Lutenegger 2007 Mase amp Hashiguchi 2009) A
direct approach based on full-size footings provides a more reliable means for foundation evaluation
The compiled database is given in Table B3 with a total of 70 large footings and plates These include a
listing of 34 foundations on 13 sands 11 footings on 4 silt deposits 13 footings or large plate load tests
on 6 intact clays and 12 foundations on 5 fissured clays Most of the footings were square or nearly
square (80) while the remaining were circular The largest foundation was a mat 5 m by 14 m in plan
loaded by stacking Kentledge concrete blocks to cause a bearing failure in the underlying soft clay For
sands the largest footing consisted of the 6-m square concrete pad In terms of equivalent square
footings mean footing sizes were 155 m (sands) 114 m (silts) and 126 m (clays)
B-16
Additional details on the individual load tests and site conditions are given elsewhere (Mayne 2009
Mayne amp Illingworth 2010 Uzielli amp Mayne 2011 2012 Mayne et al 2012 Mayne amp Woeller 2014)
Table B3 Summary of large footings soil conditions and reference sources of database
Sand Site Location Soil Conditions Footings Numbers
Shapes and Sizes
ReferencesSource
College Station
Texas Pleistocene sand 5 Square 1 15 25 33 m Briaud amp Gibbens (1999)
Kolbyttemon Sweden Glaciofluvial sand 4 Rect B = 06 12 17 24 m
Bergdahl et al (1985 1986)
Fittija Sweden Glaciofluvial sand 3 Rect B = 06 17 24 m Bergdahl et al (1984 1985)
Alvin West Texas Alluvial sand 2 Circular D = 235 m Tand et al (1994)
Alvin East Texas Alluvial sand 2 Circular D = 22 m Tand et al (1994)
Perth Australia Silceous dune sand 4 Square B = 05 and 10 m
Lehane (2008)
Grabo T1C Sweden Compacted sand fill
1 Square B = 046 m Long (1993)
Grabo T2C Sweden Compacted sand fill
1 Square B = 063 m Long (1993)
Grabo T3C Sweden Compacted sand fill
1 Square B = 080 m Long (1993)
Labenne France Dune sand 6 Square B = 07 and 10 m
Amar et al (1998)
Green Cove Florida Brown silty sand 1 Circular D = 182 m Anderson et al (2006)
Durbin South Africa
White fine sand 1 Square B = 609 m Kantley (1965)
Porto Portugal Residual clayey sands
1 Circular D = 12 m and 1 plate
Viana da Fonseca (2003)
Jossigny France Soft clayey silt 2 Square B = 1 m Amar et al (1998)
Tornhill Sweden Glacial Baltic till 3 Square B = 05 1 2 m Larsson (2001)
Vagverkel Sweden Stiff medium silt 3 Square B = 05 1 2 m Larsson (1997)
Vattahammar Sweden Brn-Gry layered silt 3 Square B = 05 1 2 m Larsson (1997)
B-17
Table B3 Continued
Clay Site Location Soil Conditions Footings Numbers Shapes and Sizes
ReferencesSource
Baytown Texas Fissured Beaumont 2 plates with d = 076 m Stuedlein amp Holtz (2008)
Belfast Ireland Soft clay silt ldquosleechrdquo
1 Square Pad B = 20 m Lehane (Geot Engr 2003)
Bothkennar Scotland Soft silty clay 2 Square Pads B = 22 24 m
Jardine et al (1995)
Bangkok Thailand Soft to stiff clay 4 Square 067 075 090 10 m
Brand et al (1972)
Haga Norway Stiff OC clay 2 Square Footings B = 10 m
Andersen amp Stenhammar (1982)
Rio Grande Brazil Sandy residual clay 3 Square 04 07 and 10 m
Consoli et al (1998)
Shellhaven England Soft estuarine clay Rectangular 5 m by 14 m Schnaid et al (1993)
Texas City A Texas Coast Fissured Beaumont 3 circular plates d = 058 m
Tand et al (1986)
Texas City B1 Texas Coast Fissured Beaumont 2 circular plates d = 058 m
Tand et al (1986)
Texas City B2 Texas Coast Fissured Beaumont 1 circular plate d = 058 m
Tand et al (1986)
Alvin Texas Texas Coast Fissured Beaumont 3 circular plates d = 058 m
Tand et al (1986)
For the series of footings on sands Figure B13 shows the summary of measure formation factors (rs) for
the 34 foundation load tests Also indicated on the graph are the corresponding mean qc values of the
sands averaged over 15middotB beneath the bearing elevations
Note also in sands that the qt and net resistance (qnet) are all very close to the measured qc since (a)
porewater pressure corrections for the a-net value are negligible in clean sands and (b) overburden
stresses are typically only 1 to 3 of the magnitude of qc This is especially true of shallow depths
Therefore for clean sands qnet is approximately qt which is approximately qc
B-18
Figure B13 Summary of sand formation factors with corresponding CPT resistances
As suggested by Briaud (2007) various in-situ measurements can be used to normalize the results from
cone penetration tests in this case Herein the qc from the CPT soundings in the sands are used to
normalize the footing stress axis as presented in Figure B14 It is evident that the results of the load-
displacement response of all 34 footings on 13 different sands can be captured by the characteristic
stress versus normalized displacement with the inclusion of the site-specific cone tip resistance In this
case the mean trend for all footings and sand sites indicates a simple expression shown in Equation B4
119954119943119952119952119957119946119951119944 119956 Clean quartz-silica sands = 120782 120787120790120787radic B4
119954119940 119913
B-19
Figure B14 Normalized foundation stresses vs square root of normalized displacement for
spread footings on sand (modified after Mayne et al 2012)
The results of measured force-displacement curves (Q vs s) from footing load tests are nonlinear and
thus the definition of capacity must be addressed Kulhawy (2004) identified three main curve types
that occur during load testing depicted in Figure B15 The most common response (Type A) shows no
clear peak and is observed in sands and silts as well as slow loading of insensitive clays and fissured
clays In this case ldquocapacityrdquo can be defined by the Eurocode corresponding to the load (or stress) when
sB=10 (Amar et al 1998) For foundations situated on many clays a plateau capacity is reached as
depicted as Type B response In rare cases where sensitive clays or structured soils are encountered a
Type C relationship occurs whereby the load-displacement curve reaches a peak followed by strain
softening In the one case observed in this study for Type C loading (Haga clay) the corresponding
ldquopeak capacityrdquo was reached at a pseudo-strain of sB = 4 as shown in Figure B16 followed by some
strain softening In the cases of soft clays at Belfast and Bothkennar a plunging type (plateau failure)
was achieved at around sB=7
B-20
Figure B15 Definitions of capacity from three types of observed foundation load test
responses (modified after Kulhawy 2004)
B-21
Figure B16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footings
In the case of fine-grained soils that develop excess porewater pressure during cone penetration the
use of the net total cone resistance (qtnet) which equals the cone tip resistance minus the total stress
must be considered (Lunne et al 1997) As noted earlier the correction of CPT data in sands is minimal
because of the low value of penetration porewater pressures and the fact that total overburden stress is
small relative to the cone resistance especially for shallow soundings Thus the use of qtnet may be
substituted for qc into the trend of Figure B1
B16 Undrained and Drained Capacity
The footings on sands typically show drained response with no excess porewater pressures developed
during loading For the foundations situated on silts analyses by Larsson (1997 2001) concluded that
these too show drained conditions For shallow foundations on clays the entire range of drainage cases
are possible ranging from undrained to partially-drained to fully-drained depending upon the applied
rate of loading relative to the permeability of the geomaterial If sufficient data were available for each
of the clay sites then the recent evaluation of drainage condition could be assessed using normalized
velocity criteria (Chung et al 2006) However this was not possible with the current data set Instead
B-22
the test loadings of footings on clays have essentially been assumed to primarily occur under undrained
conditions following recommendations by Tand et al (1986) For this case a limiting bearing stress can
be calculated from bearing capacity analyses and related directly to the CPT readings
From classical bearing capacity calculations using limit plasticity solutions the ultimate foundation stress
is shown in Equation B5
= 119925119940 ∙ 119956119958 B5 119954119958119949119957
where Nc equals the bearing factor for constant volume (514 for strip and 614 for square or circular
plan) and su equals the undrained shear strength of the clay For the case of soft to firm clays of low to
medium sensitivity the operational strength can be related to the preconsolidation stress (Mesri 1975
Jamiolkowski et al 1985) as shown in Equation B6
prime 119956119958 = 120782 120784120784120648119953 B6
Finally the effective preconsolidation stress has been linked directly to the net cone resistance (Chen amp
Mayne 1996 Demers amp Leroueil 2002) as shown in Equation B7
prime 120648119953 = 120782 120785120785(119954119957 minus 120648119959119952) B7
Combining Equations (B5) (B6) and (B7) results in direct expressions for undrained bearing capacity on
intact clays from CPT results as shown in Equations B8 and B8-2
Strip footing 119954119958119949119957 = 120782 120785120789120785(119954119957 minus 120648119959119952) B8
B-23
Squarecircle 119902119906119897119905 = 0445(119902119905 minus 120590119907119900) B8-2
Equation (B8-2) is in excellent agreement with Tand et al (1986) database study for the case of intact
clays and footings with no embedment Their study also provided a lower relationship for foundations
situated on jointed clays specifically recommending that the bearing stress not exceed 030 qtnet
Review of the available data in Figure 3 shows this to be rather conservative and a more realistic value
may be taken at 040 qtnet for fissured clays
B161 Normalized Undrained Footing Response
The characteristic stress vs normalized displacement curves for footings on clays exhibiting undrained
conditions can be verified using a relatively recent case study on Beaumont clay by Stuedlein amp Holtz
(2010) The Baytown Texas site was investigated using an exploration program of soil borings
piezocone soundings and laboratory testing The field testing included a large square footing (B = 276
m) and two small circular plates (D = 076 m) The measured load-displacement responses for all three
foundations are presented in Figure B17 Of course the large footing carried considerably higher loads
than the two small plates on the same soil deposit Yet using the aforementioned characteristic stress
(q) vs square root of normalized displacement (sB) essentially gave a unique relationship for all 3
foundations as presented in Figure B18 In this case utilizing the form of Equation B2 the
characteristic coefficient rs is 177 MPa for this deltaic clay formation
B-24
Figure B17 Measured load vs displacement curves for 3 shallow foundations on clay at
Baytown Texas (Stuedlein amp Holtz 2010)
B-25
Figure B18 Characteristic stress vs square root of normalized displacement for all three
foundations at Baytown site (Stuedlein amp Holtz 2010)
B17 General Direct CPT Method
In a manner analogous to the normalization of applied foundation stresses shown in Figure B14 for
footings on sands a generalized direct CPT method for shallow foundations on different soils can be
made in Equation B9
120782120787 119956 119954 = 119945119956 ∙ 119954119957119951119942119957 ∙ (119913
) lt 119954119940119938119953119938119940119946119957119962 B9
where qcapacity is equal to the foundation bearing capacity of the ground and hs is equal to the empirical
fitting term that depends on soil type Specifically hs equals 058 for sands 112 for silts 147 for
fissured fine-grained soils and 270 for clays A summary of the normalized stress versus the normalized
displacement curves for these four soil types is presented in Figure B19 where the data fitting to obtain
the hs parameter on clays are limited to (sB) lt 004 corresponding to undrained loading The data on
B-26
silts and sands are considered fully drained whereas the fissured clay subset may be partially drained to
undrained The statistical measures for obtaining the fitted parameter hs are quite good as shown in
Figure B20 with the coefficient of determinations (r2) values of 0947 for sands 088 for silts 0935 for
fissured and 0925 for intact clays Additional statistical evaluations on the database are provided by
Uzielli amp Mayne (2011)
Figure B19 Normalized foundation stress vs square root of normalized displacements
B-27
Figure B20 Applied foundation stress vs CPT-calculated stress for 67 footings
The same data are shown on Figure B21 in a log-log plot to emphasize the early parts of the load-
displacement responses of these footings The best fit line from regression analyses shows excellent
statistical correlations As detailed earlier the undrained bearing capacity of square or circular footings
on clays can be taken as approximately 040 times qtnet For silts and sands that experience drained
loading with no excess porewater pressures being developed and no clear peak value for capacity the
(sB) =10 criterion can be used In this case Equation 7 can be used directly to obtain stress q equal to
qcapacity when (sB)05 is 0316
An alternate approach is presented in Figure B22 summarizing the direct CPT method for estimating the
load-displacement-capacity of shallow square footings on four soil types The maximum values of soil
bearing capacity (qmax) can be capped at stresses corresponding to a percentage of the average
measured qtnet determined over the depth interval from the foundation bearing elevation to 15middotB
below the foundation In such an approach the foundation bearing capacity can be taken as
(qcapacityqtnet) of 02 for sands 035 for silts 040 for fissured and 045 for intact clays Using the assigned
values of the hs parameter the corresponding cut-off for the general stress-normalized displacement
relationship given by Equation 7 can be established specifically giving corresponding values of the
B-28
normalized displacements which can also be considered as pseudo-strains equal to (sB) max of 4 for
clays 7 for fissured clays 10 for silts and 12 for sands
Figure B21 Applied footing stress vs CPT calculated stress on logarithmic scales
B171 CPT Soil Material Index Ic
The soil behavioral type can be assessed indirectly by CPT using a material index (Ic) as defined by
Robertson (2009a b) shown in Equation B10
119920119940 = radic[120785 120786120789 minus 119949119952119944119928119957119951]120784 + [120783 120784120784 + 119949119952119944119917119955]120784 B10
B-29
Figure B22 Summary graph and equations of direct CPT method for evaluating the vertical
where Qtn equals the stress-normalized cone tip resistance and Fr equals the normalized sleeve friction
calculated from the cone penetrometer readings as shown in Equations B11 and B12 respectably
(119954119957minus120648119959119952)120648119938119957119950 = 119951 B11 119928119957119951 prime (120648119959119952120648119938119957119950)
119943119956 119917119955() = 120783120782120782 ∙ B12 (119954119957minus120648119959119952)
where σatm equals a reference stress equal to atmospheric pressure (σatm = 1 atm asymp 1 bar asymp 100 kPa) In
the initial evaluation the exponent n is set to n = 1 to find the soil behavioral type (SBT) based on a 9-
zonal chart as shown in Figure 23 The value of exponent n varies with soil type ranging from n = 1 in
intact clays and decreasing with increasing grain size to around approximately 075 in silts and
B-30
approximately 05 plusmn 02 in clean sands The appropriate value of exponent n is found by iteration until
conversion using the relationship (Robertson 2009b) shown in Equation B13
prime 120648119959119952 119951 = 120782 120785120790120783 ∙ 119920119940 + 120782 120782120787 ( ) minus 120782 120783120787 le 120783 120782 B13 120648119938119957119950
As the distinction that footings on intact clays subjected to fast loading are behaving primarily under
undrained conditions (ie constant volume) Robertson (2009a) suggested that this occurs when Ic gt 27
while in contrast Ic lt 25 corresponds more or less to drained loading (ie no excess porewater
pressures)
The CPT material index can be used to identify soil type and the corresponding characteristic hs value for
estimating footing load-displacement response In a number of the case studies investigated the results
of the sleeve friction readings were also available to allow the calculation of Ic with depth at these sites
Figure B23 shows the tentative relationship between the values of the hs parameter plotted versus the
corresponding CPT material index with an approximate trend given by Equation B14
120784120785 119945119956 = 120784 120790 minus 120783120787 B14
119920119940 120783+( ) 120784120786
B-31
Figure B23 Foundation soil formation parameter hs versus CPT material index Ic
Soil behavioral type from the Ic ranges given by Robertson (2009b) and are shown in Figure B23 with an
overall general agreement with the known soil classifications In summary the CPT can provide an
evaluation of the load-displacement-capacity response directly via Equation B15 which may be of
interest in mixed soil types such as silty sands sandy silts and the like
120784120785 Footing stress 119954 = 119954119951119942119957 ∙ radic(119956frasl119913) ∙ [120784 120790 minus
)120783120787] B15
120783+(119920119940frasl120784120786
As discussed earlier foundation bearing capacity may be taken by a limiting value of pseudo-strain (sB)
max or by qmax as specified as a percentage of qtnet This definition of bearing capacity can be seen in
comparison to soil type as seen in Figure B24
B-32
Figure B24 Bearing capacity ratio qmaxqnet versus CPT material index Ic
B172 Rectangular Foundations
This same elastic solution from Giroud (1968) for a CPT method for square and circular foundations were
utilized for rectangular foundations The study covered rectangular distortions (AB) ranging from 1
(square) to very long foundations with AB equal to 20 where A represents the foundation length and B
represents the foundation width (Mayne amp Dasenbrock 2017) The influence factor for rectangular
shaped foundations is given in Figure B25 and the expression in Equation B16
= (119912119913)120782120785120786120787 119920119912119913 B16
B-33
Figure B25 Influence factor for rectangular foundations from elastic theory solution (Mayne
amp Dasenbrock 2017)
Including 32 full scale load tests on sands results from an additional 98 shallow foundations have been
reported in previous studies The foundations were primarily square or circular and among these include
large spread footings with measured settlements and bearing capacity Table B4 shows a summary of
the number of footings and footing sizes used in the study The range of length to width ratios varied
from 1 lt (AB) lt 23 with an average value of (AB) equal to 238 The embedment depth to footing
width ranged from 0 lt DeB lt 222 and averaged 046
Table B4 Sources for shallow foundation performace
Source of Data No of
Footings
Mean B
(m)
Max B
(m)
Min B
(m)
Mean A
(m)
Max A
(m)
Min A
(m)
Mayne et al (2012) 32 149 609 046 149 609 046
Lehane (2011) 8 041 027 060 041 060 027
Gifford et al (1987) 17 846 1588 527 1591 3596 701
Jeyapalan amp Boehm
(1986)
13 1022 2892 400 1671 3049 640
B-34
Schmertmann
(1970)
31 945 5608 061 1288 8668 061
Papadopoulos
(1992)
29 870 3600 100 1300 7290 100
Total 130 671 5608 027 1012 8668 027
The solution for shallow foundation settlements given back in Equation B1 combined with
the expression in Equation B16 takes on the form
119954∙119913∙119920119912119913∙(120783minus120642120784) 119956 = B17
119916119956
Combining the direct CPT equation developed from the 32 full scale load tests (Equation B15)
together with the influence factor for rectangular shaped foundations becomes
minus120782120785120786120787 120784120785 119912 Footing stress 119954 = 119954119951119942119957 ∙ radic(119956frasl119913) ∙ [120784 120790 minus
)120783120787] ∙ [ ] B18
120783+(119920119940frasl120784120786 119913
B18 Conclusions
A direct CPT method for square rectangular and circular shallow footings is developed using a database
of 166 full-scale field load tests where the minimum footing size of an equivalent square width of 05 m
is established to avoid scaling issues The use of load vs displacement is generalized by characteristic
stress vs normalized displacement (sB) and further simplified by a square root plotting technique that
captures the ground response in a single parameter that is related to the qtnet Data are grouped
according to four main soil categories sands silts fissured clays and intact clays It is generally believed
that the footings on sands and silts exhibit fully drained behavior while in intact clays an undrained
response occurs under conditions of constant volume
B-35
The summary equation for evaluating the vertical stress-displacement-capacity of square rectangular
and circular footings is given by
minus120782120785120786120787 119956 119912 119954119943119952119952119957119946119951119944 = 119945119956 ∙ 119954119957119951119942119957 ∙ radic ∙ ( ) lt 119954119950119938119961 B19
119913 119913
where the parameter hs also relates directly to Ic The foundation capacity depends upon the mode of
failure as well as drainage conditions and can be taken as a fraction of qtnet where qmaxqtnet is 020 in
sands 035 in silts 040 in fissured clays and 045 in intact clays as shown in Figure 24 Alternatively the
capacity can be defined using a limiting pseudo-strain given by (sB) max of 4 in clays 7 in fissured
10 in silts and 12 in sands
B-36
APPENDIX C
List of Figures
Figure C1 Components of Axial Pile Capacity C-3
Figure C2 Selected alpha relationships as function of the undrained shear strengthC-5
Figure C3 Pile side friction coefficient β in terms of effective friction angle and overconsolidation ratio
C-11
Figure C4 Concept of Direct versus Rational Method for CPT evaluation of axial pile capacityC-14
Figure C5 Soil behavior type using CPT via original UniConeC-19
Figure C6 Soil behavior type using CPT via Modified UniCone C-19
Figure C7 Nine-zone soil behavioral soil type using normalized piezocone parameters C-21
Figure C8 Summary of Modified UniCone Method for direct CPT assessment of axial pile capacity C-22
Figure C9 Elastic continuum solution for axial pile response under compression and tension loadingC-24
List of Tables
Table C1 Alpha Methods for Axial Pile Side Friction where fp = αmiddotsu C-5
Table C2 Beta Methods for Axial Pile Side Friction where fp = βmiddotσvoC-9
Table C3 Selection of Direct CPT Methods for Axial Pile CapacityC-15
C-1
C1 EVALUATING DEEP FOUNDATION RESPONSE FROM CONE
PENETRATION TESTS
C11 Introduction
The axial response of deep foundations includes the load-displacement-capacity and axial load transfer
when driven pilings and drilled shafts are loaded in compression and uplift The use of cone penetration
testing (CPT) especially seismic piezocone tests (SCPTu) are advantageous since they provide
information on the subsurface soils and their geomaterial properties The SCPTu provides at least four
measurements on soil behavior with depth including (a) cone tip resistance qt (b) sleeve friction fs (c)
penetration porewater pressure u2 and (d) shear wave velocity Vs This offers the opportunity to
determine the geostratigraphy unit weight effective overburden stress shear strength stress state
and stiffness of the ground from a single exploratory sounding thus values in economic expediency
and reliability The axial pile capacity can be calculated based on static equilibrium of forces acting along
the sides and base of the pile foundation The displacement of the pile can be ascertained using
elasticity theory from closed-form solutions boundary elements andor finite element analyses Load
transfer occurs along the length of the pile and only a portion of axial forces are transmitted to the base
or toe or tip of the pile These too can be assessed using elasticity solutions
An alternate approach to pile capacity involves the utilization of direct CPT methods available
since the 1970s but in the past 10+ years a number of new and statistically reliable algorithms
have been developed which can be implemented for highway design and construction
C111 Axial Pile Capacity
The axial compression capacity of a single pile foundation is composed of a shaft or side component and
end-bearing component at the base as depicted in Figure 1 For a circular pile the side capacity (Qs) is
determined from the unit side friction (fp) acting along the surface area of the shaft which is As = πmiddotdmiddotL
where d = pile diameter and L = length embedded below grade
If the magnitude of side friction is uniform and constant with depth the side capacity is simply
C-2
119928119956 = 119943119953 middot 119912119956 C1
Moreover however many piles extend through multiple layers and a summation of unit side
frictions acting on various pile segments must be tabulated over the length of the pile as
suggested by Figure C1
Figure C1 Components of Axial Pile Capacity
The unit-end bearing resistance (qb) acts over the base of the pile tip where the area of a circular pile is
given by Ab = πmiddotd24 For piles in compression loading the base capacity is determined from Equation
C2
119876119887 = 119902119887 middot 119860119887 C2
and for piles in tension (or uplift) it is normally taken that Qb = 0
C-3
C112 Pile Unit Side Friction
Several different approaches can be adopted for evaluating the unit pile side friction (fp) prior to full-
scale load testing and construction (Poulos amp Davis 1980 ONeill 2001) The most common types
including the alpha and beta methods as summarized in Tables 1 and 2 respectively In the alpha
method an empirical coefficient (α) is applied to the undrained shear strength (su) of clay soils to
obtain
119891119901 = 120572 middot 119904119906 C3
An illustrative example of alpha curves is shown in Figure C2 A difficulty with the alpha
method is the evolution of many variants and changes to the expressions for its estimation It is
also restricted in its use for specific types of piles in clays and fine-grained soils
C-4
Figure C2 Selected alpha relationships as function of the undrained shear strength
Table C1 Alpha Methods for Axial Pile Side Friction where fp = αmiddotsu
C-5
Reference
Source
Alpha Equation Remarks
Tomlinson
(1957)
2716801102
uu ssAll pile types (steel
concrete timber)
Note su in ksf
Tomlinson
(1957)
2616201102
uu ssConcrete piles
Note su in ksf
McClelland
(1974)
Kerisel α = 07 - 031middotln(su)
Peck α = 10 - 02middotsu
Woodward α = 091middot(su)091
Driven piles in clay
Note su in ksf
Semple
(1980) )1(511
)1(1
OCR
OCR Summary from 9 series of
pile load tests
American
Petroleum
Institute (API
1981)
For suσvo le 035 α = 10
For suσvo gt 080 α = 05
Otherwise α = 1 + 1111middot (035 - suσvo)
Driven steel pipe piles
Tomlinson
(1986)
1 α = 55 (su)-091
2 α = 785 (su)-102
3 α = 605 (su)-100
where 75 lt su (kPa) lt 210
1 Driven piles
2 Bored piles
3 Driven piles in till with L
lt 10middotd
API (1987) 1 α = 10 for su lt 25 kPa Driven piles in clay other
than Gulf of Mexico
C-6
2 α = 125 - 001middotsu for 25 lt su lt 75 kPa
3 α = 050 for su gt 75 kPa
Kulhawy amp
Jackson
(1989)
α = 021+026middot(σatm su) le 1 Analyses of 106 drilled
shaft foundations
Table C1 Continued
API (1989) 05 For suσ vo le 1 α =
su radic frasl prime σvo
minus025 su For suσ vo gt 1 α = 05 ∙ ( frasl prime ) σvo
Driven steel pipe piles
Kulhawy amp
Jackson
(1989)
OCR
OCR
sin50
tan)sin1( sin
Λ = (1-CsCc) asymp 080 for
insensitive clays
Nowacki et al
(1996)
minus05 su For suσvo le 07 α = 05 ∙ ( frasl prime ) σvo
minus02 su For suσvo gt 07 α = 055 ∙ ( frasl prime ) σvo
Driven pile foundations
Kolk and van
der Velde
(1996)
α =09middot FL middot (suσvo)-03 lt 10
FL = [ (L - z)d]-02 = length term
L = pile length
Note su obtained from
UU lab tests
Applicable to driven piles
in clays
C-7
Knappett amp α = 1 for su le 30 kPa Non-displacement piles in
Craig (2012) fine-grained soils α = 116 - (su185) for 30 kPa le su le 150 kPa
α = 035 for su gt 150 kPa
Knappett amp
Craig (2012) α = 055 ∙ 02 40
( ) (LfraslD)
∙ su ( frasl prime σvo
minus03
) Displacement piles in fine-
grained soils
z = depth at point considered
d = outside diameter of pile
Karlsrud et al
(2005 NGI
Method)
α = function (suσvo and plasticity index) Driven pilings
Fleming et al
(2009) 05 minus05 su su For su σvo le 1 120572 = ( frasl prime ) ∙ ( frasl prime )
σvo NC σvo
05 minus025 su su For su σvo gt 1 α = ( frasl prime ) ∙ ( frasl prime ) σvo NC σvo
For driven piles in clay
where (suσvo)NC is the
normalized shear strength
for normally-consolidated
clay
Brown et al
(2010)
α = 0 for 0 lt z le 5 feet
α = 055 for z gt 5 feet and (suσatm) le 15
α = 055-01middot(suσatm -15) for 15le (suσatm) le 25
Drilled Shafts and Bored
Piles
Table C1 Continued
C-8
OCRK )sin1(0
Karlsrud
(2012)
α = function (su σvo and plasticity index) Driven pilings
Note as reported by Karlsrud (2012)
In the beta method the coefficient β is applied to the effective overburden stress (σvo) at the point of
concern along the pile length
prime = 120573 C4 119891119901 middot 120590119907119900
Table C2 Beta Methods for Axial Pile Side Friction where fp = βmiddotσvo
Reference Source Beta Equation Remarks
Burland (1973) β = (1-sinϕ) middot tan ϕ Piles in NC clays
Meyerhof (1976) β = Kmiddottanδ where K = K0 = for NC clays
K = 15middotK0 for OC clays
Poulos amp Davis
(1980)
β = (1-sinϕ) middot tan ϕ middot OCR05 Piles in OC stiff clays
Table C2 Continued
C-9
Kulhawy amp Jackson
(1989)
β = (1-sinϕ) middot tan ϕ middot OCRsinϕ Piles in quartz sands and insensitive
clays
ONeill (2001) β = (1-sinϕ) middot tan θ middot OCRsinϕ Drilled shafts where θ = (δϕ)middotϕ and
δ = interface friction between pile and soil
Fleming et al
(2009)
β = Kmiddottanδ K = lateral stress coefficient and δ =
friction angle between soil and pile
material
Karlsrud (2012) β = fctn (OCR and PI) PI = plasticity index of the clay ()
Mayne and Niazi
(2017)
β = CM middot CK middot K0 middot tanϕ
where K0 = (1-sinϕ) middot OCRsinϕ for soils
that are virgin loaded then unloaded
CM = pile material factor = 1 (drilled
augered) 09 (prestressed or precast
concrete) 08 (timber) and 07
(steel) and CK = pile installation factor
= 09 (bored or augered) 10 (low
displacement eg H-pile or open-end
pipe) and 11 (driven high
displacement eg prestressed
concrete closed-end pipe)
As the beta approach applies to all types of soils (gravels sands silts and clays) and more or
less has remained unchanged since its advent circa 1970 it has been selected for further
discussion herein In the direct form for calculation of unit pile side friction the expression is
given by
119874119862119877119904119894119899120601 prime prime 119891119901 = 119862119872 ∙ 119862119870 ∙ (1 minus 119904119894119899120601prime) middot ∙ 119905119886119899120601prime ∙ 120590119907119900 C5
C-10
where CM is a soil-pile interface friction coefficient and CK = pile installation factor Values of CM are in
the range 07 lt CM lt 10 depending upon pile material (steel wood concrete) and ranges for CK vary
09 lt CK lt 11 depending upon method of installation (auger drill driven) as detailed in Table 2
The value of K0 is limited to the passive stress coefficient which for the simple Rankine case is given by
KP = (1+sin ϕ) (1 - sin ϕ ) Thus there is a maximum value of overconsolidation ratio for which
Equation C5 applies given by OCRlimit = [(1+sin ϕ ) (1 - sin ϕ )2] (1sinϕ)
The corresponding graph in Figure C3 shows the relationship for β in terms of ϕ and OCR
Figure C3 Pile side friction coefficient β in terms of effective friction angle and overconsolidation ratio
C113 Pile Unit End Bearing
C1131 Theoretical Considerations
The evaluation of the end bearing resistance of pile foundations is commonly determined using
limit plasticity theory where
prime Drained loading = C6 119902119887 119873119902 middot 120590119907119900
C-11
Undrained loading 119902119887 = 119873119888 middot 119904119906 C7
where the value of σvo is calculated at the depth z = L and the value of su is the average undrained shear
strength from z = L to a depth z = L + d beneath the pile tip For a circular pile the limit plasticity solution
for undrained loading gives (Vesic 1977) a value Nc = 933 while a deep strip foundation would employ a
value of Nc = 824 For drained loading the expression for Nq for a deep foundation is given by (Vesic
1977)
)]arctan()sin1(tan21[)](tan1[sin1
sin1)tanexp( 2 BLABNq
C8
where A and B are the pile plan dimensions (for a circular pile A = B) and L = pile length For a
circular pile this can be approximated by
asymp 077(120601prime75deg) 119873119902 C9
over a range of effective friction angles 20deg le ϕ le 45deg
C1132 Practical Considerations
For drained loading of sands the full calculated end bearing capacity will never be realized because it
would require the pile to move a distance equal to its diameter This is beyond practical use and
therefore the limit plasticity solutions must be clipped to a fraction of the calculated value Another
reason for using a reduced end-bearing capacity is due to strain incompatibility since the side capacity is
mobilized early while the end-bearing is engaged much later So to have compatible values of Qs and
Qb the qb must be reduced using the following guidelines (Randolph 2003 Mayne 2007)
prime prime C10 119902119887 = 119873119902 ∙ 120590119907119900 ∙ 119891119909
where fx = strain incompatibility factor
fx = 010 for drilled shafts augered cast-in-place and bored piles
fx = 020 for driven low displacement piles (opened-ended steel pipe and H-piles)
fx = 030 for driven high-displacement piles (ie solid piles PSC and closed-ended pipe)
C-12
For undrained loading of circular piles in compression no reduction of end-bearing is necessary
therefore
119902119887 = 933 middot 119904119906 C11
C12 Direct CPT Methods for Pile Capacity
In the direct CPT method the penetrometer readings are scaled directly via specified algorithms to
obtain the pile unit side friction and end-bearing as depicted in Figure C4 As many as 40 different
direct CPT methods have been developed over the past five decades as summarized by Niazi amp Mayne
(2013) Starting circa 1970 many of these early methods relied on hand-recorded data from field
mechanical CPTs where only qc data were obtained at 20 cm intervals or later with mechanical readings
of both qc and fs using special sets of inner and outer rods that recorded vertical load for tip and loads
for tip plus sleeve in alternating increments Also early pile load test data were often obtained from
top-down measurements of load-displacement
C-13
Undrained qb = Nc su
Qside = S (fp DAs)
QTotal = Qs + Qb - Wp
fp = cmck Ko svorsquo tanrsquo
Qbase = qb Ab
Drained qb = Nq svorsquo
Method Two Rational or ldquoIndirectrdquo Method
Method OneldquoDirectrdquo CPT Method
(Scaled Pile)
fp = fctn (soil type pile
type qt fs and u2)
qb = fctn (soil type qt-u2)
qb = unit end bearing
unit sidefriction fp
OCR su Ko t DR rsquo
AXIAL PILE CAPACITY FROM CPT READINGS
Figure C4 Concept of Direct versus Rational Method for CPT evaluation of axial pile capacity
Beginning in the mid-1990s as the modern electric piezocone (CPTu) was implemented the newer
equipment offered improved data and better resolution because of the additional reading of porewater
pressures and correction of raw measured qc to total resistance qt In addition the use of electronic
digital data collection and field computers proved superior in field measurements and recordings As a
consequence several reliable CPT methods for axial pile capacity have been developed Moreover
parallel improvements in full-scale pile load testing occurred and now include modern strain gage
instrumentation digital data recording and automated testing procedures Also the testing can be
single direction (compression or tension) or bi-directional as in the Osterberg cell
C-14
Table C3 provides a selection of recent direct CPT methods that have become available over the
past two decades A number of these (namely ICP NGI UWA and Fugro) were funded by the
offshore industry because of the growth of oil amp gas reserves and windfarm installations thus
necessitating increased concerns on risk probability and reliability in the site investigations for
offshore platforms and design of large driven monopile foundations These direct CPT methods
are often statistically based on large datasets compiled from full-scale load tests made
worldwide (eg Schneider et al 2008)
Table C3 Selection of Direct CPT Methods for Axial Pile Capacity
Method Pile Types Soil Types References CPT
data
Additional
parameters
needed
Unicone Method all types Sands silts
clays
Eslami and
Fellenius
(1997)
Fellenius
(2009)
qt fs u2
KTRI = Kajima
Technical
Research
Institute
all types Sands
mixed clays
Takesue et al
(1998)
fs and u2 No guidance given
on end bearing
resistance
Imperial College
Procedure (ICP)
OE and CE Sands Chow et al
(1997 PhD)
Jardine et al
(2005)
qt Interface friction
angle (δ) from
ring shear tests
Correlated to
mean grain size
(D50)
C-15
OE and CE Clays Jardine et al
(2005)
qt Interface friction
angle (δ)
correlated to PI
Table C3 Continued
NGI Method
(Norwegian
Geotechnical
Institute)
OE and CE Sands Clausen et al
(2005)
qt DR from qt
Clays a Alpha
method
(Karlsrud et al
2005 2012)
b Beta
method
(Karlsrud
2012)
qt for su
qt for
OCR
PI = plasticity
index ()
PI = plasticity
index ()
Fugro Method OE and CE Sands Kolk et al
(2005)
qt
Clays Van Dijk and
Kolk (2011)
qt
OE and CE Sands Lehane et al
(2005)
qt IFR = infilling ratio
related to amount
of plugging
C-16
Purdue LFRD Drilled
shafts
Sands Basu amp
Salgado
(2012)
qt LRFD = load
resistance
factored design
Enhanced Various Sands silts Niazi and qt u2 Based on 330 load
Unicone Method clays and Mayne (2015 and fs tests including
mixed soils 2016) bored augered
jacked and driven
piles
UWA (Univ of
Western
Ra = pile
roughness
Australia) Clays Lehane et al
(2012)
qt Interface friction
angle (δ) from
ring shear tests
and also method
without δ
HKU Method
(Hong Kong
Univ)
OE and CE
(end
resistance
only)
Sands Yu and Yang
(2012)
qt End bearing only
PLR = plug length
ratio
Note PLR can be
estimated from
OE inner diam
Table C3 Continued
Note previously called Marine Technical Directorate (MTD)
C-17
Many of the offshore CPT methods relate to driven piles either in sand or clay as that is
common deep foundation for those purposes Of particular interest are the Unicone and
Modified Unicone Methods since they use all three readings of the piezocone (qt fs and u2)
and address a variety of pile foundation types
C13 Modified UniCone Method
The modified UniCone Method was developed based on a total 330 pile load tests which were
associated with SCPTu data during their site investigations (Niazi and Mayne 2015 2016) This
represents a threefold increase over the original UniCone database that was built upon data
from 106 pile load tests (Eslami amp Fellenius 1997)
For the original UniCone algorithms use is made of the effective cone resistance (qE)
119902119864 = 119902119905 minus 1199062 C12
and a chart of qE vs fs provided an approximate soil classification in five distinct groups as shown by
Figure C5 Later in the modified approach a better delineation of the larger dataset gave soil
subclassifications as indicated by Figure C6
C-18
Figure C5 Soil behavior type using CPT via original UniCone
Figure C6 Soil behavior type using CPT via Modified UniCone
C-19
In the modified approach the 9-zone normalized soil behavioral type (SBTn) is ascertained using CPT
data in conjunction with the Robertson (2009) charts as shown in Figure C7 As discussed in Appendix
A and B the SBTn method uses the normalized cone resistance (Qtn) normalized sleeve friction (Fr())
and CPT material index Ic This permits a much wider range in the calculated pile side friction because fp
is related as a continuous curve with Ic rather than only 5 values that are assigned in the original
scheme
The pile unit side friction (fp) is obtained from qE and the CPT material index Ic using the following
expression at each elevation along the sides of the pile
10(0732 middot 119868119888 minus 3605) fp middot middot C13 = 119902119864 middot 120579119875119879 middot 120579119879119862 120579119877119860119879119864
C-20
Figure C7 Nine-zone soil behavioral soil type using normalized piezocone parameters
(after Robertson 2009 Mayne 2014)
where θPT = coefficient for pile type (θPT = 084 for bored piles 102 for jacked piles 113 for driven
piles) θTC = coefficient for loading direction (θTC = 111 for compression and 085 for tension) and θRATE
= rate coefficient applied to soils in SBT zones 1 through 7 (θRATE = 109 for constant rate of penetration
tests and 097 for maintained load tests) Since CPT provides data at regular intervals of 2 cm to 5 cm
along the sides of the pile the average fp from z = 0 to z = L can be used directly in Equation 1 to obtain
the shaft capacity Qs
C-21
The pile end bearing resistance is obtained from
middot 10(0325middot 119868119888minus1218) 119902119887 = 119902119864 C14
where qE is averaged in the vicinity of the pile tip Figure C8 shows the Modified Unicone Method in
graphical format For sensitive clays of zone 1 please see additional discussions by Niazi amp Mayne
(2016)
Figure C8 Summary of Modified UniCone Method for direct CPT assessment of axial pile
capacity
C-22
C14 Axial Pile Displacement
The movement of pile foundations can be assessed using elastic continuum theory (Poulos amp
Davis 1980 Randolph 2003 Mayne amp Niazi 2017) These relationships have been developed
using finite element analyses boundary elements and analytical closed-form solutions For the
latter approach the top displacement of a rigid pile subjected to a vertical force is shown in
Figure C9 This gives the movement at the top of a pile subjected to either compression or
tension (uplift) loading Also the percentage of axial load transferred to the pile toe is
determined For piles extending through various soil layers the elastic solution can be
implemented by using a set of stacked pile segments each with its own stiffness as
represented by a soil Youngs modulus
For pile groups the use of computer software is recommended Several available programs can handle
pile groups under axial and lateral moment loading such as DEFPIG (Univ Sydney) and PIGLET (Univ
Western Australia) A full listing of pile foundation software is given at the Geotechnical amp
GeoEnvironmental Service Directory wwwggsdcom
Ps
= Sh
aft
Load
sL
t
tEd
IPw
Load Transfer to Base
)]1)((5ln[
)(
)1(1
1
2 vdL
dLFI
E
ECT
211
IF
P
P CT
t
b
Base Load = Pb
Randolph Elastic Solution
Ground Surface
Top Load Pt = Ps + Pb
Rigid Pile Response
wt = displacementd = diameter
L = length
Es = Elastic Soil Modulus
Es0 (at z = 0) surface
EsM (at z = L2)mid-length
EsL (at z = L)full length
Load Transfer to Shaft
E = EsMEsL = Gibson parameterE = 1 for homogeneous case (constant E)E = 05 for pure Gibson case (Es0 = 0)
where FCT = load direction (= 1 compression 0 tension)
21
IF
P
P CT
t
b
z = depth
= Poissons ratio
Tension
Compression
C-23
Figure C9 Elastic continuum solution for axial pile response under compression and tension
loading
C-24
C15 Acknowledgements
The author appreciates the help of Fernando Illingworth of PonsEcuador Meena Viswanth of
GeoSyntecAtlanta and Dr Marco Uzielli of GeoRiskItaly in helping to collect and process the data
Funding support was provided by David Woeller of ConeTec Richmond BC
C-25
- Structure Bookmarks
-
- CONE PENETRATION TEST DESIGN GUIDE FOR STATE GEOTECHNICAL ENGINEERS
- TABLE OF CONTENTS
- LIST OF TABLES
- CHAPTER 1 INTRODUCTION
- CHAPTER 2 DIRECT CPT METHOD FOR SOIL CHARACTERIZATION
- CHAPTER 3 DIRECT CPT METHOD FOR SHALLOW FOUNDATIONS
- CHAPTER 4 DIRECT CPT METHOD FOR DEEP FOUNDATIONS
- REFERENCES
- APPENDIX A
- APPENDIX B
- APPENDIX C
-
TABLE OF CONTENTS
CHAPTER 1 INTRODUCTION1
CHAPTER 2 DIRECT CPT METHOD FOR SOIL CHARACTERIZATION 4
21 Introduction 4
22 Soil Unit Weight 6
23 CPT Material Index 7
231 Step 1 Normalized Sleeve Friction 7
232 Step 2 Iteration 8
24 Soil Behavior Type (SBT) 8
241 Step 1 8
242 Step 2 9
25 Effective Stress Friction Angle 9
26 Stress History 10
27 Lateral Stress Coefficient 11
28 Undrained Shear Strength 12
29 Ground Stiffness and Soil Moduli 12
210 Coefficient of Consolidation 13
211 Hydraulic Conductivity14
212 Example Problems 15
2121 Example 1 Direct CPT Methods for Geoparameters on Sands 15
2122 Example 2 Direct CPT Methods for Geoparameters on Clay 33
CHAPTER 3 DIRECT CPT METHOD FOR SHALLOW FOUNDATIONS 56
31 Procedure 56
311 Step 1 Estimating Footing Dimensions57
312 Step 2 Soil Characterization 58
313 Step 2a Foundation soil formation parameter59
314 Step 3 Soil elastic modulus and Poissonrsquos ratio 60
315 Step 4 Net cone tip resistance 60
316 Step 5 Bearing capacity of the soil 61
317 Step 6 Settlement61
318 Step 7 Final Check 61
32 Example Problems 62
321 Example 3 Direct CPT Method on Sands 62
CHAPTER 4 DIRECT CPT METHOD FOR DEEP FOUNDATIONS 70
41 Introduction70
42 Modified UniCone Method72
421 Step 172
422 Step 272
423 Step 373
424 Step 473
43 Axial Pile Displacements 73
44 Example Problems 74
441 Example 5 Direct CPT Methods Axial Pile Capacity74
REFERENCES 87
APPENDIX A
APPENDIX B
APPENDIX C
1
2
3
4
5
6
7
8
9
10
11
12
13
14
15
16
17
18
19
20
21
22
23
24
LIST OF FIGURES
Figure Geoparameters determined from CPT 1
Figure Conventional method for shallow foundation design compared to direct CPT method 2
Figure Direct CPT evaluation of axial pile capacity 3
Figure Soil unit weight from CPT sleeve friction 6
Figure Approximate ldquorules of thumbrdquo method using CPT sounding from Wakota Bridge MN 7
Figure SBT zones using CPT Ic 9
Figure Yield stress exponent compared to CPT material index 11
Figure k vs dissipation time for 50 consolidation (Mayne 2017) 15
Figure CPT data from Benton County Minnesota for example problem 1 16
Figure Soil layers using ldquorules of thumbrdquo17
Figure Soil layer 1 using SBT method 23
Figure Soil layer 2 using SBT method 24
Figure Soil layer 3 using SBT method 25
Figure Soil layer 4 using SBT method 26
Figure CPT data from Minnesota for example problem 2 34
Figure Soil layers using ldquorules of thumbrdquo36
Figure Soil layer 1 using SBT method 42
Figure Soil layer 2 using SBT method 43
Figure Soil layer 3 using SBT method 44
Figure Soil layer 4 using SBT method 45
Figure Dissipation t50 data 53
Figure Direct CPT method for shallow foundations56
Figure Direct CPT method introduction 58
Figure Differentiation of porewater pressure measurement locations (Lunne et al 1997) 58
Figure 25 Foundation soil formation parameter hs versus CPT material index Ic (Mayne 2017) 60
Figure 26 Diagram of footing profiles for Example 362
Figure 27 CPT data from Northern Minnesota 63
Figure 28 Schematic of foundation design associated with CPT data 64
Figure 29 Soil type for example problem 367
Figure 30 Direct CPT evaluation of axial pile capacity 70
Figure 31 UniCone Method soil behavior type using CPT (Mayne 2017) 71
Figure 32 Modified UniCone Method soil behavior type using CPT (Mayne 2017) 71
Figure 33 Deep foundation end bearing pile diagram74
Figure 34 CPT data from Minnesota for example problem 5 75
Figure 35 Soil layers using ldquorules of thumbrdquo for pile capacity example using direct CPT method 76
Figure 36 Soil layer 1 using SBT method 81
Figure 37 Soil layer 2 using SBT method 82
Figure 38 Soil layer 3 using SBT method 83
Figure 39 Soil layering compared to pilings 85
LIST OF TABLES
Table 1 Geoparameters calculated directly from CPT 4
Table 2 Minor Geoparameters 5
Table 3 Soil type compared to exponent m 10
Table 4 Geoparameters evaluated for Case Example 117
Table 5 Geoparameters 35
Table 6 Bearing capacity defined by soil type61
Table 7 SCPT Results 62
CHAPTER 1 INTRODUCTION
The purpose of this manual is to provide an analysis of soil characterization shallow footings and deep
foundations using direct cone penetration testing (CPT) methods Geotechnical site characterization is
important for evaluating soil parameters that will be used in the analysis and design of foundations
retaining walls embankments and situations involving slope stability Common practice is to determine
these soil parameters through conventional lab and in-situ testing An alternative method uses CPT
readings of cone tip resistance (qt) sleeve friction ( fs) and pore pressure (u2) directly to determine these
parameters such as unit weight effective friction angle undrained shear strength and many others
(Figure 1) Direct CPT methods are provided for these parameters which will be used in the designs for
shallow and deep foundations
Figure 1 Geoparameters determined from CPT
Designing shallow foundations is typically done in a two-part process determining the bearing capacity
and expected settlement (commonly referred to as displacement) of the soil to approximate the
required size and shape of a foundation The older traditional methods are no longer required with
many approaches existing for using CPT directly in the design of shallow foundations Results from these
methods can provide a direct assessment of bearing capacity andor settlement A specific approach to
the direct method has been recommended and tested using a database of 166 full-scale field load tests
(Figure 2) A one-part process is used to scale the measured cone penetrometer readings (ie
measured cone tip resistance sleeve friction and pore water pressure) to obtain the bearing capacity
and settlement of the soil A step-by-step procedure has been created to transition from the CPT data to
bearing capacity with settlement accounted for The steps consist of estimating a design footing width
1
and length while using the process in the Soil Characterization section to determine soil parameters
directly from the CPT data
Figure 2 Conventional method for shallow foundation design compared to direct CPT method
There are upwards of 40 different direct CPT methods that have been developed over the past five
decades to determine the axial compression capacity of a piling foundation Earlier direct CPT methods
relied on hand-recorded information where mechanical-type CPT cone tip resistance data would be
collected at 20 cm intervals
Herein the method recommended for deep foundation design is the Modified UniCone method which
uses all three readings of the modern electronic piezocone penetrometer (CPTu) while addressing a
variety of pile foundation types (Figure 3) The modified UniCone Method is based on a total of 330 pile
load tests (three times the original UniCone database) that were associated with SCPTu data Two
computer software programs have been recommended for analyzing pile movements andor
settlements
2
Figure 3 Direct CPT evaluation of axial pile capacity
3
Symbol Parameter
γt Soil total unit weight
Ic CPT material index
SBT Soil behavior type (SBT)
ʹ σp Preconsolidation stress
YSR Yield stress ratio
ϕʹ Effective friction angle
Ko Lateral stress coefficient
su Undrained shear strength
Dʹ Constrained modulus
Eʹ Drained Youngrsquos modulus
CHAPTER 2 DIRECT CPT METHOD FOR SOIL
CHARACTERIZATION
21 INTRODUCTION
A direct CPT method for determining the value of each of various geoparameters shown in Table 1 is
provided The parameter Ic is used in the derivation of several sequential parameters This section of
the guide may be referred to as these parameters are used in calculations throughout the shallow
foundations and deep foundations sections
Table 1 Geoparameters calculated directly from CPT
4
Kʹ Bulk modulus
ks Subgrade reaction modulus
MR Resilient modulus
Gmax Small-strain shear modulus
k Coefficient of permeability
cv Coefficient of consolidation
Symbol Parameter Equation
ρt Mass Density ρt = γtga
where ga = 98 ms2
σvo Total Stress σvo asymp Σ (γti middot Δzi)
c Effective cohesion ʹ Empirical c asymp 003σp
In clays c asymp 01cu
Other minor parameters can be used in calculations of the geoparameters in Table 1 These minor parameters are provided in Table 2
Table 2 Minor Geoparameters
5
22 SOIL UNIT WEIGHT
The total soil unit weight can be estimated from CPT sleeve friction resistance as shown in Figure 4
(Mayne 2014) This method is not applicable to organic clays diatomaceous soils peats or sensitive
soils
Figure 4 Soil unit weight from CPT sleeve friction
Use Equation 1 to calculate the soils total unit weight
1 120574119905 = 120574119908 ∙ [122 + 015 ∙ ln (100 ∙ 119891119904 + 001)]
120590119886119905119898
Since soil unit weight is required for determining most geoparameters (including Soil Behavior Type)
estimating the soil type of the layers using the ldquorules of thumbrdquo method is a good first step before
determining more precise layering (Figure 5) Once a unit weight is determined for each noticeable layer
change these results can be used in later calculations such as ldquoCPT Material Indexrdquo To use the ldquorules
of thumbrdquo method some helpful guidelines are to assume sands are identified when qt gt 725 psi and u2
asymp uo while the presence of intact clays are prevalent when qt lt 725 psi and u2 gt uo The magnitude of
porewater pressures help to indicate intact clays such as soft (u2 asymp 2middotuo) firm (u2 asymp 4middotuo) stiff (u2 asymp 8middotuo)
and hard (u2 asymp 20middotuo) Fissured overconsolidated clays tend to have negative u2 values such that u2 lt 0
6
Figure 5 Approximate ldquorules of thumbrdquo method using CPT sounding from Wakota Bridge MN
23 CPT MATERIAL INDEX
The development of the CPT material index Ic has improved the initial classification of soil types and
calculation of soil parameters as shown in Table 1 To calculate Ic follow steps 1 and 2
231 Step 1 Normalized Sleeve Friction
Calculate Fr using Equation 2 if not provided with the CPT data gathered
2
7
119891119904 119865119903() = 100 ∙
(119902119905minus120590119907119900)
232 Step 2 Iteration
Iterate using Equations 3-5 to determine Ic by initially using n = 1 to calculate a starting value of Ic The
exponent n is soil-type dependent n = 1 (clays) n asymp 075 (silts) and n asymp 05 (sands) Iteration converges
quickly which is generally after the 3rd cycle
3 (119954119957minus120648119959119952)120648119938119957119950 = 119951 119928119957119951 prime (120648119959119952120648119938119957119950)
4
prime 120590119907119900 119899 = 0381 ∙ 119868119888 + 005 ( ) minus 015 120590119886119905119898
119899 le 10
5 119920119940 = radic[120785 120786120789 minus 119949119952119944119928119957119951]120784 + [120783 120784120784 + 119949119952119944119917119955]120784
24 SOIL BEHAVIOR TYPE (SBT)
Typically soil samples are not taken when using CPT The soil types are then inferred from the qt fs and
u2 readings To determine the types of soil from CPT data follow steps 1 and 2
241 Step 1
To determine the soil layers from CPT results calculate Ic by following the steps under section ldquoCPT
Material Indexrdquo After Ic has been determined through all specified depths use Figure 6 to classify the
type of soil by comparing each Ic value to normalized CPT readings (Fr and Qtn) from Equations 2 and 3
8
Figure 6 SBT zones using CPT Ic
242 Step 2
To determine if any of the soil layers contain ldquosensitive clays and siltsrdquo from zone 1 or ldquovery stiff
overconsolidated (OC) soilrdquo from zones 8 and 9 use Equations 6 and 7 If any soil layers are found within
zone 1 by Equation 6 then caution should be taken as these clays are prone to instability collapses and
difficulties in construction performance Very stiff OC sands to clayey sands of zone 8 (15 lt Fr lt 45)
and very stiff OC clays to silts of zone 9 (Fr gt 45) can be identified by Equation 7
6 Equation 6 errata This is an exponential expression (see Figure 5)
119876119905119899 lt 12119890119909119901(minus14 ∙ 119865119903)
7 119876119905119899 gt 0005(119865119903minus1)minus00003(119865119903minus1)2minus0002
1
Errata See terms and coefficients in Figure 5 above (numbers cannot be rounded off)
After each CPT reading has been assigned a zone from Figure 14 a visual representation can be made to
show the predominant layers by soil types (Figure A5)
25 EFFECTIVE STRESS FRICTION ANGLE
The effective friction angle (ϕ) is used to govern the strength for sands and clays where Equation 8 is
used for sands and Equation 9 is used for clays The value of ϕ for sands is derived from Ic so before ϕ
can be calculated refer to the iteration of Qtn under the ldquoCPT Material Indexrdquo section Once the values
9
Soil Type m
Fissured clays 11
Intact clays 10
Sensitive clays 09
Silt mixtures 085
Silty sands 080
Clean sands 072
Note m may be higher than 11 in fissured clays
of Ic and Qtn are calculated the type of soil can be determined from the section ldquoSoil Behavior Type
SBTrdquo The type of soil will dictate which equation to use for ϕ
Sands
8 120601prime(deg) = 176deg + 110deg log(119876119905119899)
Clays
9 119902119905minus120590119907119900 120601prime(deg) = 295deg ∙ 119861119902
0121 ∙ [0256 + 0336 ∙ 119861119902 + log ( )] prime 120590119907119900
Where = 119861119902 (1199062 minus 119906119900)(119902119905 minus 120590119907119900)
26 STRESS HISTORY
Determining the stress history can be characterized by an apparent yield stress ratio of the form
10 prime 120590119901
119884119878119877 = prime 120590119907119900
Where σp is defined as the preconsolidation stress or effective yield stress (Equation 10) The YSR is the
same equation as the more common overconsolidation ratio (OCR) but is now generalized to
accommodate mechanisms of preconsolidation such as ageing desiccation repeated cycles of wetting-
drying repeated freeze thaw cycles and other factors
11 prime prime 120590119901 = 120590119910 = 033(119902119905 minus 120590119907119900)119898prime
Where σp σy σvo and qt have units of kPa and the value of m depends on soil type with typical values shown in Table 3
Table 3 Soil type compared to exponent m
10
The value of m for non-fissured soils and inorganic clays and silts is derived from Ic (Figure 7) so before
m can be calculated (Equation 12) refer to the iteration of Ic under the ldquoCPT Material Indexrdquo section
Determine Ic for all soil layers before calculating m
12 028
119898prime = 1 minus 25 119868119888 1+( frasl ) 265
Once m is known for all soil layers the YSR can be determined
Figure 7 Yield stress exponent compared to CPT material index
A limiting value of YSR can be reached for clays and sands It can be calculated using Equation 13
13 (1 sin(emptyprime))
(1+sin(emptyprime)) 119884119878119877119897119894119898119894119905 = [ ]
(1minussin(emptyprime))2
27 LATERAL STRESS COEFFICIENT
The lateral stress coefficient Ko = σhoσvo commonly referred to as the at-rest condition is used to
represent the horizontal geostatic state of soil stress Ko can be calculated using Equation 14 but
11
14
119870119900 = (1 minus sin(emptyprime)) ∙ 119884119878119877sin( emptyprime)
sections such as ldquoStress Historyrdquo and ldquoEffective Stress Friction Anglerdquo will need to be referred to for
determining parameters ϕ and YSR
A maximum value for Ko can be determined by Equation 15
15 (1+sin(emptyprime))
119870119900119898119886119909 = = 1199051198861198992(45deg + emptyprime2) (1minussin(emptyprime))
28 UNDRAINED SHEAR STRENGTH
Loading on soils can result in fully drained partially drained or fully undrained conditions Sands
typically produce drained cases due to their high permeability but exceptions may occur in loose sands
during fast loading where the water does not have sufficient time to dissipate Clays exhibit low
permeability and thus often result in undrained loading cases when a load is applied quickly For soft-
firm clays the undrained shear strength (su) can be determined from CPT via Equation 16 where the
value of the bearing factor Nkt can be taken as 12
16 119902119905minus120590119907119900 119904119906 =
119873119896119905
In the case of remolded undrained shear strength from CPT su asymp fs
29 GROUND STIFFNESS AND SOIL MODULI
Determining the grounds stiffness can be measured from geoparameters such as the constrained
modulus (D) drained Youngrsquos modulus (E) bulk modulus (K) subgrade reaction modulus (ks) resilient
modulus (MR) and small-strain shear modulus (Gmax)
17 119863 prime asymp 5 ∙ (119902119905 minus 120590119907119900)
18
12
119864 prime = 119863 prime
11
19 119864prime
119870prime = [3∙(1minus2119907prime)]
The subgrade modulus (ks) is a combination of soil-structural properties which creates a parameter that
depends on the ground stiffness and the size of the loaded element
20 119864prime
119896119904 = [119889∙(1minus1199072)]
The resilient modulus MR applies to pavement analysis and design and can be calculated using Equation
21 where qt and fs are in MPa
21 053 119872119877 = (146119902119905 + 135511989111990414 + 236)244
The small strain shear modulus Gmax is a representation of the initial stiffness of all soils and rocks
Graphically it is the beginning portion of all stress-strain-strength curves for geomaterials Use Equation
22 to determine Gmax
22 119866119898119886119909 = 120588119905 ∙ 1198811199042
Where Vs = shear wave velocity (Equation 23) as measured by seismic cone penetration tests (SCPT) If
only cone penetration tests (CPT) or piezocone (CPTu) data are available the shear wave velocity may
be estimated from
23 03 119891119904 119881119904 (119898frasl119904) = [101 ∙ log(119902119905) minus 114]167 ∙ (100 ∙ )
119902119905
Where qt and fs have units of (kPa)
210 COEFFICIENT OF CONSOLIDATION
The coefficient of consolidation (cv) controls the rate that foundation and embankment settlements
occur By using results of CPT dissipation tests that measure the change in u2 readings over time the
value of cv can be determined Using Equation 24 cv can be determined based on piezocone dissipation
13
curves The equation below requires an estimate of the in-situ rigidity index (IR) of the soil If results of
SCPTU are available then IR may be determined from Gmax and qt per equation A38
24 0030∙(119886119888)2∙(119868119877)075
119888119907 = 11990550
Where ac = penetrometer radius (178 cm for 10-cm2 cone 220 cm for 15-cm2 cone) t50 = time to reach 50 dissipation IR = Gsu = undrained rigidity index G can be determined from the ldquoGround Stiffness and Soil Modulirdquo section
211 HYDRAULIC CONDUCTIVITY
The hydraulic conductivity also known as the coefficient of permeability (k) expresses the flow
characteristics of soils and has units of cms or feetday One method of calculation would be to use
Equation 25 where cv would need to be determined from ldquoCoefficient of Consolidationrdquo and D would
need to be determined from ldquoGround Stiffness and Soil Modulirdquo
25 119888119907∙120574119908 119896 =
119863prime
An alternative approach developed for soft normally-consolidated soils (Figure 8) is shown in Equation
26 where t50 (sec) values are used directly in assessing k in (cms)
26
14
125
119896 asymp ( 1
) 251∙11990550
Figure 8 k vs dissipation time for 50 consolidation (Mayne 2017)
212 EXAMPLE PROBLEMS
2121 Example 1 Direct CPT Methods for Geoparameters on Sands
Several geoparameters need to be determined based on the given CPT data collected for the South
Abutment of a bridge in Benton County MN (Figure 9) The groundwater table (GWT) was measured at
17 feet Determine all the geoparameters found in Table 4 at depths of 0 feet to 30 feet All sand layers
can be assumed ldquodrainedrdquo with ν = 02
15
Figure 9 CPT data from Benton County Minnesota for example problem 1
16
Table 4 Geoparameters evaluated for Case Example 1
Symbol Parameter Symbol Parameter
γt Soil total unit weight Ko Lateral stress coefficient
Ic CPT material index Dʹ Constrained modulus
SBT Soil behavior type (SBT) Eʹ Drained Youngrsquos modulus ʹ σp Preconsolidation stress Kʹ Bulk modulus
YSR Yield stress ratio MR Resilient modulus
ϕʹ Effective friction angle Gmax Small-strain shear modulus
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 10 Soil layers using ldquorules of thumbrdquo
γw = 6224 pcf
17
Layer 1
From Figure 10 fs = 17 psi taken as a representative value of the layer
17 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1204 pcf
145 psi
Layer 2
From Figure 10 fs = 17 psi taken as a representative value of the layer
17psi
γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1204 pcf 145 psi
Layer 3
From Figure 10 fs = 12 psi taken as a representative value of the layer
12 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 4
From Figure 10 fs = 7 psi taken as a representative value of the layer
CPT Material Index
7 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1121 pcf
145 psi
Layer 1
From Figure 10 fs = 17 psi and qc = 3500 psi
18
qt = qc + u2 ∙ (1 minus a) = 3500 psi + 3 psi ∙ (1 minus 08) = 3501 psi
= γt ∙ 6 feet = 1204 pcf ∙ 6 feet = 7225 psf = 50 psi σvo
fs 17 psi Fr() = 100 ∙ = 100 ∙ = 049
(qt minus σvo) (3501 psi minus 50 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 5 psi minus 0 psi = 50 psi
(qt minus σvo)σatm (3501 psi minus 50 psi )145 psi = = Qtn prime )n (σvoσatm (50 psi145 psi)n
prime σvo 50 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(049)]2
Qtn = 35329 n = 036 lt 10 Ic = 13
Layer 2
From Figure 10 fs = 17 psi and qc = 3500 psi
19
qt = qc + u2 ∙ (1 minus a) = 3500 psi + 0 psi ∙ (1 minus 08) = 3500 psi
= 7225 psf + γt ∙ 6 feet = 7225 psf + 1204 pcf ∙ 6 feet = 1445 psf = 100 psi σvo
fs 17 psi Fr() = 100 ∙ = 100 ∙ = 049
(qt minus σvo) (3500 psi minus 100 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 100 psi minus 0 psi = 100 psi
(qt minus σvo)σatm (3500 psi minus 100 psi )145 psi = = Qtn prime )n (σvoσatm (100 psi145 psi)n
prime σvo 100 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
47 minus log(Qtn)]2 + [122 + log(049)]2
Ic = radic[3
Qtn = 27947 n = 041 lt 10 Ic = 14
Layer 3
From Figure 10 fs = 12 psi and qc = 1500 psi
20
qt = qc + u2 ∙ (1 minus a) = 1500 psi + 0 psi ∙ (1 minus 08) = 1500 psi
= 1445 psf + γt ∙ 11 feet = 1445 psf + 1172 pcf ∙ 11 feet = 2734 psf = 190 psi σvo
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 081
(qt minus σvo) (1500 psi minus 190 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = γwater ∙ (z minus 119911119908) = 6224 pcf ∙ (23 feet minus 17 feet) = 3734 psf = 99 psi
prime σvo = σvo minus uo = 190 psi minus 99 psi = 164 psi
(qt minus σvo)σatm (1500 psi minus 190 psi )145 psi = = Qtn prime )n (σvoσatm (164 psi145 psi)n
prime σvo 164 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(081)]2
Qtn = 947 n = 06 lt 10 Ic = 19
21
Layer 4
From Figure 10 fs = 7 psi and qc = 1200 psi
qt = qc + u2 ∙ (1 minus a) = 1200 psi + 0 psi ∙ (1 minus 08) = 1200 psi
= 2734 psf + γt ∙ 11 feet = 2734 psf + 1121 pcf ∙ 7 feet = 3519 psf = 244 psi σvo
fs 7 psi Fr() = 100 ∙ = 100 ∙ = 060
(qt minus σvo) (1200 psi minus 244 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = γwater ∙ (z minus 119911119908) = 6224 pcf ∙ (30 feet minus 17 feet) = 8091 psf = 130 psi
prime σvo = σvo minus uo = 244 psi minus 130 psi = 188 psi
(qt minus σvo)σatm (1200 psi minus 244 psi )145 psi = = Qtn prime )n (σvoσatm (188 psi145 psi)n
prime σvo 188 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(060)]2
22
Qtn = 686 n = 064 lt 10 Ic = 19
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic Qtn and Fr the first layer is defined as a ldquoDrained Gravelly Sandrdquo from
Figure 11
Figure 11 Soil layer 1 using SBT method
23
Layer 2
Based on values of Ic Qtn and Fr the second layer is defined as a ldquoDrained Sandrdquo from Figure
12
Figure 12 Soil layer 2 using SBT method
24
Layer 3
Based on values of Ic Qtn and Fr the third layer is defined as a ldquoDrained Sandrdquo from Figure 13
Figure 13 Soil layer 3 using SBT method
25
Layer 4
Based on values of Ic Qtn and Fr the fourth layer is defined as a ldquoDrained Sandrdquo from Figure 14
Figure 14 Soil layer 4 using SBT method
Effective Stress Friction Angle
Layer 1
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(3533) = 456deg
Layer 2
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(2795) = 445deg
26
Layer 3
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(947) = 393deg
Layer 4
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(686) = 378deg
Stress History
Layer 1
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (13frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(24139 kPa minus 345 kPa)072 = 4717 kPa
= 684 psi
prime σp 684 psi YSR = = = 136
primeσvo 50 psi
Layer 2
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + (14 1 + ( frasl ) frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(34132 kPa minus 689 kPa)072 = 605 kPa
= 877 psi
prime σp 877 psi YSR = = = 87
primeσvo 100 psi
27
Layer 3
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + (19 1 + ( frasl ) frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(10342 kPa minus 131 kPa)072 = 254 kPa
= 368 psi
prime σp 368 psi YSR = = = 22
primeσvo 164 psi
Layer 4
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (19frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(8274 kPa minus 168 kPa)072 = 215 kPa = 312 psi
prime σp 312 psi YSR = = = 17
primeσvo 188 psi
Lateral Stress Coefficient
Layer 1
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(456deg)) ∙ 136sin( 456deg) = 18
Layer 2
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(445deg)) ∙ 87sin( 445deg) = 14
28
Dprime 17480 psi
Eprime = = = 15890 psi 11 11
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(393deg)) ∙ 22sin( 393deg) = 06
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(378deg)) ∙ 17sin( 378deg) = 05
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3501 psi minus 50 psi) = 17480 psi
Eprime 15890 psi
Kprime = = = 8828 psi [3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Layer 3
Layer 4
Ground Stiffness and Soil Moduli
Layer 1
Values of qt and fs need to be in MPa
MR = (146(241 MPa)053 + 1355(012 MPa)14 + 236)244 = 3416 MPa = 49575 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1172 kPa m Vs (mfrasls) = [101 ∙ log(24139 kPa) minus 114]167 ∙ (100 ∙ ) = 2745
24139 kPa s
Vs = 9012 fts
29
γt 1204 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 37 ∙ (9012 ) = 304 ∙ 106psf = 21000 psi s
Layer 2
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3500 psi minus 100 psi) = 17480 psi
Dprime 17480 psi Eprime = = = 15890 psi
11 11
Eprime 15890 psi Kprime = = = 8828 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(241 MPa)053 + 1355(012 MPa)14 + 236)244 = 3415 MPa = 49562 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1172 kPa m Vs (mfrasls) = [101 ∙ log(24133 kPa) minus 114]167 ∙ (100 ∙ ) = 2745
24133 kPa s
Vs = 9012 fts
γt 1204 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
30
2 ft Gmax = ρt ∙ Vs
2 = 37 ∙ (9012 ) = 304 ∙ 106psf = 21000psi s
Layer 3
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (1500 psi minus 190 psi) = 7405 psi
Dprime 7405 psi Eprime = = = 6732 psi
11 11
Eprime 6732 psi Kprime = = = 3740 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(103 MPa)053 + 1355(008 MPa)14 + 236)244 = 1506 MPa = 21854 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 827 kPa m Vs (mfrasls) = [101 ∙ log(10343 kPa) minus 114]167 ∙ (100 ∙ ) = 2611
10343 kPa s
Vs = 8566 fts
γt 1172 pcf slug ρt = = = 36
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 36 ∙ (8566 ) = 27 ∙ 106psf = 19000 psi s
Layer 4
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (1200 psi minus 244 psi) = 5878 psi
31
Dprime 5878 psi Eprime = = = 5343 psi
11 11
Eprime 5343 psi Kprime = = = 2969 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(83 MPa)053 + 1355(005 MPa)14 + 236)244 = 165 MPa = 16901 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 483 kPa m Vs (mfrasls) = [101 ∙ log(8274 kPa) minus 114]167 ∙ (100 ∙ ) = 2243
8274 kPa s
Vs = 7360 fts
γt 1121 pcf slug ρt = = = 35
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 35 ∙ (7360 ) = 389 ∙ 106psf = 13000 psi s
32
2122 Example 2 Direct CPT Methods for Geoparameters on Clay
Several geoparameters need to be determined based on the given CPT data collected for the South
Abutment (Figure 15) The groundwater table (GWT) was measured at 60 feet Determine all the
geoparameters found in Table 5 at depths of 0 feet to 42 feet All sand layers can be assumed ldquodrainedrdquo
ν = 02 All clay layers can assume ν = 049
33
Figure 15 CPT data from Minnesota for example problem 2
34
Table 5 Geoparameters
Symbol Parameter Symbol Parameter
γt Soil total unit weight Eʹ Drained Youngrsquos modulus
Ic CPT material index Kʹ Bulk modulus
SBT Soil behavior type (SBT) MR Resilient modulus
ʹ σp Preconsolidation stress Gmax Small-strain shear modulus
YSR Yield stress ratio su Undrained shear strength
ϕʹ Effective friction angle cv Coefficient of consolidation
Ko Lateral stress coefficient k Hydraulic conductivity
Dʹ Constrained modulus
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
35
Figure 16 Soil layers using ldquorules of thumbrdquo
Unit weight of water γw = 6224 pcf
Layer 1
From Figure 16 fs = 13 psi taken as a representative value of the layer
13 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1179 pcf
145 psi
Layer 2
From Figure 16 fs = 12 psi taken as a representative value of the layer
12 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 3
From Figure 16 fs = 20 psi taken as a representative value of the layer
36
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
Layer 4
From Figure 16 fs = 2 psi taken as a representative value of the layer
2 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1004 pcf
145 psi
CPT Material Index
Layer 1
From Figure 10 fs = 13 psi and qc = 3000 psi
qt = qc + u2 ∙ (1 minus a) = 3000 psi + 0 psi ∙ (1 minus 08) = 3000 psi
σvo = γt ∙ 2 feet = 1179 pcf ∙ 2 feet = 235 psf = 16 psi
fs 13 psi Fr() = 100 ∙ = 100 ∙ = 043
(qt minus σvo) (3000 psi minus 16 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 16 psi minus 0 psi = 16 psi
(qt minus σvo)σatm (3000 psi minus 16 psi )145 psi = = Qtn prime )n (16 psi145 psi)n (σvoσatm
prime σvo 16 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(043)]2
37
Qtn = 4126 n = 032 lt 10 Ic = 121 therefore sand
Layer 2
From Figure 10 fs = 12 psi and qc = 250 psi
qt = qc + u2 ∙ (1 minus a) = 250 psi + 10 psi ∙ (1 minus 08) = 252 psi
σvo = 235 psf + γt ∙ 30 feet = 235 psf + 1172 pcf ∙ 30 feet = 3751 psf = 260 psi
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 53
(q minus σ ) (252 psi minus 260 psi) t vo
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 260 psi minus 0 psi = 260 psi
(qt minus σvo)σatm (250 psi minus 260 psi )145 psi = = Qtn prime (σvoσatm)n (260 psi145 psi)n
38
prime σvo 260 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(53)]2
Qtn = 87 n = 10 le 10 Ic = 32 therefore clay
Layer 3
From Figure 10 fs = 20 psi and qc = 4000 psi
qt = qc + u2 ∙ (1 minus a) = 4000 psi + 8 psi ∙ (1 minus 08) = 40016 psi
σvo = 3751 psf + γt ∙ 2 feet = 3751 psf + 1219 pcf ∙ 2 feet = 3995 psf = 277 psi
fs 16 psi Fr() = 100 ∙ = 100 ∙ = 054
(qt minus σvo) (30016 psi minus 277 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 277 psi minus 0 psi = 277 psi
(qt minus σvo)σatm (30016 minus 277)145 psi = = Qtn prime (σvoσatm)n (277 psi145 psi)n
39
primeσvo 277 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(054)]2
Qtn = 1962 n = 052 le 10 Ic = 15 therefore sand
Layer 4
From Figure 10 fs = 2 psi and qc = 250 psi
qt = qc + u2 ∙ (1 minus a) = 250 psi + 0 psi ∙ (1 minus 08) = 250 psi
= 3994 psf + γt ∙ 8 feet = 3994 psf + 1004 pcf ∙ 8 feet = 4798 psf = 333 psi σvo
fs 2 psi Fr() = 100 ∙ = 100 ∙ = 092
(qt minus σvo) (250 psi minus 333 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 333 psi minus 0 psi = 333 psi
(qt minus σvo)σatm (250 psi minus 333 psi )145 psi = = Qtn prime (σvoσatm)n (333 psi145 psi)n
40
prime σvo 333 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(092)]2
Qtn = 65 n = 10 lt 10 Ic = 29 therefore clayey silt
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic Qtn and Fr the first layer is defined as a ldquoDrained Sandrdquo from Figure 17
41
Figure 17 Soil layer 1 using SBT method
Layer 2
Based on values of Ic Qtn and Fr the second layer is defined as a ldquoUndrained Clayrdquo from Figure
18
42
Figure 18 Soil layer 2 using SBT method
43
Layer 3
Based on values of Ic Qtn and Fr the third layer is defined as a ldquoDrained Sandrdquo from Figure 19
Figure 19 Soil layer 3 using SBT method
44
Layer 4
Based on values of Ic Qtn and Fr the fourth layer is defined as a ldquoUndrained Silty Mixrdquo from
Figure 20
Figure 20 Soil layer 4 using SBT method
Effective Stress Friction Angle
Layer 1 (sand)
45
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(4126) = 464deg
Layer 2 (clay)
qt minus σvo
ϕprime(deg) = 295deg ∙ Bq0121 ∙ [0256 + 0336 ∙ Bq + log ( )]
primeσvo
Where
(u2 minus uo) (10 psi minus 0 psi) Bq = = = 004
(qt minus σvo) (252 psi minus 260 psi)
252 psi minus 260 psi ϕprime(deg) = 295deg ∙ 00440121 ∙ [0256 + 0336 ∙ 0044 + log ( )]
260 psi
ϕprime(deg) = 245deg
Layer 3 (sand)
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(1962) = 428deg
Layer 4 (clay)
qt minus σvo
ϕprime(deg) = 295deg ∙ Bq0121 ∙ [0256 + 0336 ∙ Bq + log ( )]
primeσvo
Where
(u2 minus uo) (5 psi minus 0 psi) Bq = = = 002
(qt minus σvo) (251 psi minus 333 psi)
252 psi minus 333 psi ϕprime(deg) = 295deg ∙ 0020121 ∙ [0256 + 0336 ∙ 002 + log ( )]
333 psi
ϕprime(deg) = 265deg
Stress History
46
Layer 1
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (12frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(3000 psi minus 33)072 = 1051 psi
prime σp 1051 psi YSR = = = 642
primeσvo 16 psi
Layer 2
028 028
prime m = 1 minus = 1 minus = 099 25 25 1 + (
Icfrasl ) 1 + (32frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(252 psi minus 260)099 = 734 psi
prime σp 734 psi YSR = = = 28
primeσvo 260 psi
Layer 3
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + ( frasl ) 1 + (15frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(4002 psi minus 277)072 = 1288 psi
47
prime σp 1288 psi YSR = = = 46
primeσvo 277 psi
Layer 4
028 028
prime m = 1 minus = 1 minus = 097 25 25 Ic 1 + ( frasl ) 1 + (29frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(250 psi minus 333)097 = 626 psi
prime σp 626 psi YSR = = = 19
primeσvo 333 psi
Lateral Stress Coefficient
Layer 1
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(464deg)) ∙ 642sin( 464deg) = 56
Layer 2
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(245deg)) ∙ 28sin( 245deg) = 090
Layer 3
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(428deg)) ∙ 46sin( 428deg) = 091
Layer 4
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(266deg)) ∙ 19sin( 266deg) = 073
48
Ground Stiffness and Soil Moduli
Layer 1
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3000 psi minus 16 psi) = 14991 psi
Dprime 14991 psi Eprime = = = 13629 psi
11 11
Eprime 13629 psi Kprime = = = 7571 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(207 MPa)053 + 1355(009 MPa)14 + 236)244 = 2816 MPa = 40873 psi
fs
03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 896 kPa m Vs (mfrasls) = [101 ∙ log(20685 kPa) minus 114]167 ∙ (100 ∙ ) = 2564
20685 kPa s
Vs = 8412 fts
γt 1179 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
2 ft 2 Gmax = ρt ∙ Vs = 37 ∙ (8412 ) = 260 ∙ 106psf = 17992 psi
s
Layer 2
49
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (252 psi minus 260 psi) = 11298 psi
Dprime 11298 psi Eprime = = = 10271 psi
11 11
Eprime 10271 psi Kprime = = = 17118 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(049))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(17 MPa)053 + 1355(008 MPa)14 + 236)244 = 443 MPa = 64309 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 827 kPa m Vs (mfrasls) = [101 ∙ log(17375 kPa) minus 114]167 ∙ (100 ∙ ) = 2646
17375 kPa s
Vs = 8680 fts
γt 1172 pcf slug ρt = = = 36
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 36 ∙ (8680 ) = 274 ∙ 106psf = 19036 psi s
Layer 3
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (4002 psi minus 277 psi) = 19869 psi
Dprime 19869 psi Eprime = = = 18063 psi
11 11
50
Eprime 18063 psi Kprime = = = 10035 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(276 MPa)053 + 1355(014 MPa)14 + 236)244 = 4017 MPa = 58309 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1379 kPa m Vs (mfrasls) = [101 ∙ log(27591 kPa) minus 1379]167 ∙ (100 ∙ ) = 2854
27591 kPa s
Vs = 9363 fts
γt 121 pcf slug ρt = = = 38
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 38 ∙ (9363 ) = 33 ∙ 106psf = 23050 psi s
Layer 4
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (250 psi minus 333 psi) = 1088 psi
Dprime 1088 psi Eprime = = = 989 psi
11 11
Eprime 989 psi Kprime = = = 16490 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(049))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
51
MR = (146(17 MPa)053 + 1355(001 MPa)14 + 236)244 = 360 MPa = 52189 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 138 kPa m Vs (mfrasls) = [101 ∙ log(17238 kPa) minus 114]167 ∙ (100 ∙ ) = 1545
17238 kPa s
Vs = 5069 fts
γt 1004 pcf slug ρt = = = 31
ga 322 ftfrasls2 ft3
2 ft 2 Gmax = ρt ∙ Vs = 31 ∙ (5069 ) = 80 ∙ 105psf = 5565 psi
s
Undrained Shear Strength
Layer 2
qt minus σv0 250 psi minus 26 psi su = = = 187 psi
12 12
Layer 4
qt minus σv0 250 psi minus 333 psi su = = = 181 psi
12 12
Coefficient of Consolidation
52
Example dissipation data are shown in Figure 21 Example calculations are provided
Figure 21 Dissipation t50 data
Layer 1
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (7197)075
cv = = = 359 cmsec 1 sec t50
Layer 2
53
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (94392)075
cv = = = 0008 cmsec 3000 sec t50
Layer 3
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (8303)075
cv = = = 665 cmsec 06 sec t50
Layer 4
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (26714)075
cv = = = 087 cmsec 11 sec t50
Hydraulic Conductivity
Layer 1
cm 1 psi 359 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 10 ∙ 10minus4 cmsec Dprime 14984 psi
Layer 2
cm 1 psi 0008 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 73 ∙ 10minus7 cmsec Dprime 11379 psi
Layer 3
cm 1 psi 665 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 20 ∙ 10minus5 cmsec Dprime 148650 psi
54
Layer 4
cm 1 psi 087 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 18 ∙ 10minus4 cmsec Dprime 10861 psi
55
CHAPTER 3 DIRECT CPT METHOD FOR SHALLOW
FOUNDATIONS
31 PROCEDURE
Shallow foundation analysis is typically done in a two-part traditional procedure The traditional
techniques are no longer required as a direct CPT method for square rectangular and circular shallow
footings is available (Figure 22) This process has the soil types grouped into four main categories sands
silts fissured clays and intact clays When determining soil types for each design it is believed footings
on sands and silts act in a fully drained manner while intact clays act in an undrained manner under
conditions of constant volume In order to determine the vertical stress-displacement-capacity of
square rectangular and circular shallow footings follow the steps provided towards the solution given
by Equation 27
Figure 22 Direct CPT method for shallow foundations
Equation 27 may be used to calculate all footing stresses from zero to the bearing capacity (qmax) To
calculate qmax for a sized footing of width (B) length (L) and thickness (t) follow steps 1 through 7
provided The settlement (s) can be determined after the calculation of qmax by simply rearranging
56
Equation 27 where the allowable stress (qallow) is defined as qmax divided by the factor of safety (FS) For
shallow footings a FS value of 3 is commonly used in geotechnical engineering
120782120787 minus120782120785120786120787 119956 119923 119954119950119938119961 = 119945119956 ∙ 119954119957119951119942119957 ∙ (119913
) ∙ (119913
) 27 119950119938119961
2 119902119898119886119909 0345 1 frasl119865119878 119871 119904 = 119861 ∙ [ ∙ ∙ ( ) ] 28
119861 ℎ119904 119902119905119899119890119905
Where hs = the foundation soil formation parameter qtnet = the net corrected cone tip resistance
311 Step 1 Estimating Footing Dimensions
In Minnesota frost heave can have devastating effect on a shallow foundations It is common practice to
place a foundation bearing elevation below the expected maximum frost depth (roughly 45 to 6 feet
below ground level) With this assumption estimate a footing size (B x L) for design to obtain
representative data from CPT roughly 15middotB below the foundation depth (Df) as shown in Figure 23 The
CPT data collected will consist of
cone tip resistance (qc)
measured porewater pressure acting behind the cone tip (u2) as shown in Figure 24
sleeve friction (fs)
57
Figure 23 Direct CPT method introduction
Figure 24 Differentiation of porewater pressure measurement locations (Lunne et al 1997)
312 Step 2 Soil Characterization
The soil behavior type (SBT) and CPT material index (Ic) govern the value of the formation factor hs The
first step is following the steps in Soil Unit Weight to determine soil layering and γt values for each
layer To determine a representative unit weight (γsoil) for all the layers in the range of Df to 15∙B below
58
Df the user will need to use their engineering judgement on the definition of ldquorepresentative unit
weightrdquo This single value of unit weight is used in further calculations such as the total vertical soil stress (σvo) from Equation 29 and effective vertical stress (σvo) from Equation 30 Values of σvo and σvo
will need to be calculated at 15middotB below Df
120648119959119952 = sum(120632119956119952119946119949 ∙ 119963) 29
prime 120648119959119952 = 120648119959119952 minus 119958119952 30
Where z = Df +15middotB uo=γwatermiddot(z-zw)
The qc will need to be corrected using Equation 31 in the case of fine-grained soils that develop excess porewater pressure during cone penetration These values will be used in the following calculations
119954119957 = 119954119940 + 119958120784 ∙ (120783 minus 119938) 31
Where 119860119899 a = cone area ratio = eg MnDOT commonly uses a = 08 119860119888
119860119899 = cross-sectional area of load cell or shaft 119860119888 = projected area of the cone
The cone area ratio is determined based on the type of piezocone tip used during in-situ field testing
Manufacturer specifications should provide the measured net area ratio (a) for the particular cone
penetrometer as determined by calibration in a pressurized triaxial chamber
313 Step 2a Foundation soil formation parameter
With the representative value γsoil continue to follow the steps in CPT Material Index through Soil
Behavior Type (SBT) to better define the type of soil at depth 15∙B below Df Use the value of Ic
calculated at depth 15∙B below Df to calculate hs with Equation 36 This parameter is based on the soil
type with typical values shown in Figure 25 The data on silts and sands are considered fully drained
whereas the fissured clay subset may be partially drained to undrained
59
23 ℎ119904 = 28 minus
119868119888 15 32
1+( ) 24
Figure 25 Foundation soil formation parameter hs versus CPT material index Ic (Mayne 2017)
314 Step 3 Soil elastic modulus and Poissonrsquos ratio
Details about the soil such as its elastic modulus (Es) and Poissonrsquos ratio (ν) will be needed for further
calculations A representative value for the Es can be determined from in-situ field tests Values of ν can
be assigned as 02 for drained sands and as 05 for undrained loading cases involving clays (Jardine et al
1985 Burland 1989)
315 Step 4 Net cone tip resistance
Calculate qtnet the mean value of net cone tip resistance 15middotB below the foundation bearing elevation
using Equation 33
33 119902119905119899119890119905 = 119902119905 minus 120590119907119900
60
316 Step 5 Bearing capacity of the soil
Use the assumed B and L calculated hs and qtnet to determine the soils bearing capacity (qmax) from
Equation 34 Use Table 6 to calculate qmax by using the maximum allowable settlement ratio (sB)max
correlating to the soil type If hs is in between the given values interpolate to acquire (sB)max
Table 6 Bearing capacity defined by soil type
Type of Soil hs (sB)max
Clean Sands 058 12
Silts 112 10
Fissured Clays 147 7
Intact Clays 270 4
05 minus0345 119904 119871 119902119898119886119909 = ℎ119904 ∙ 119902119905119899119890119905 ∙ (119861
) ∙ (119861
) 34 119898119886119909
317 Step 6 Settlement
Settlement can be calculated directly using the results from Equation 35 A FS equal to 3 is common in
foundation engineering
2 119902119898119886119909 0345 1 frasl119865119878 119871 119904 = 119861 ∙ [ ∙ ∙ ( ) ] 35
119861 ℎ119904 119902119905119899119890119905
318 Step 7 Final Check
Check that the applied stress (q) is less than qmax using Equation 36 Repeat the process again with a
new B andor new L if q gt qmax
05 119904 ) 119902 = ℎ119904 ∙ 119902119905119899119890119905 ∙ (119861
minus0345 119871 ∙ ( )
11986136
61
32 EXAMPLE PROBLEMS
321 Example 3 Direct CPT Method on Sands
A footing size needs to be determined based on the given CPT data collected for the South Abutment
(Figure 27) The footing stress (q) was determined to be 8000 psf Estimate a footing size (B x L) and
determine the bearing capacity of the foundation (qmax) using the direct CPT method provided Also
determine the expected settlement based on the calculated bearing capacity
Figure 26 Diagram of footing profiles for Example 3
The soil elastic modulus determined from seismic CPT (SCPT) are shown in Table 7
Table 7 SCPT Results
Depth (feet) Es (tsf)
Bottom of layer
3 557
6 433
9 557
12 695
17 590
22 501
27 571
32 505
62
Figure 27 CPT data from Northern Minnesota
Solution
Step 1 Assume the footing is placed below the frost depth of 6 feet
Estimate L L = 50 feet = 600 inches
63
Estimate B B = 12 feet = 144 inches
Estimate footing thickness t t = 2 feet
Df = 6 feet = 72 inches
Df + 15middotB = 24 feet = 288 inches
Step 2 Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 28 Schematic of foundation design associated with CPT data
Layer 1
From Figure 28 fs = 20 psi taken as a representative value of the layer
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
64
Layer 2
From Figure 28 fs = 20 psi taken as a representative value of the layer
20psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
Layer 3
From Figure 28 fs = 8 psi taken as a representative value of the layer
8 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1134 pcf
145 psi
Using the unit weights calculated between Df and 15B determine a representative unit weight of the soil to calculate the total and effective soil stresses Based on layer 3 being the weakest supporting layer γsoil = 113 pcf
σvo = γsoil ∙ (119863119891 + 15 ∙ B) = 113 lbsfraslft3 ∙ 24 feet = 2721 psf = 189 psi
prime σvo = σvo minus uo = 2721 psf minus 0 psf = 2721 psf = 189 psi
Calculate the cone tip resistance between Df and Df + 15B below the foundation depth
qt = qc + u2 ∙ (1 minus a) = 1250 psi + 0 psi ∙ (1 minus 08) = 1000 psi
Step 2a CPT Material Index
Using Figure 28 the sleeve friction at 15B below the foundation depth is about 16 psi
65
fs 16 psi Fr() = 100 ∙ = 100 ∙ = 13
(qt minus σvo) (1250 psi minus 189 psi)
Iterate to solve for Ic Steps are not shown for brevity
(qt minus σvo)σatm (1250 psi minus 189 psi )145 psi = = Qtn prime (σvoσatm)n (189 psi145 psi)n
prime σvo 189 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Qtn = 703 n = 072 lt 10 Ic = 210
66
Figure 29 Soil type for example problem 3
Use the Ic value to determine the foundation soil formation parameter
Calculate the foundation soil formation parameter The tip resistance is about 1250 psi at 24 feet and
the cone area ratio was determined to be 08 from information provided by the manufacturer
23 23 hs = 28 minus = 28 minus = 078 = SandSilt 15 15 Ic 210
1 + ( ) 1 + ( ) 24 24
67
Step 3 Determine representative values of soil elastic modulus from soil testing and Poissons ratio The
soil elastic modulus was taken as the average value between depths of Df and 15middotB (6 feet and 24 feet)
433 + 557 + 695 + 590 + 501 Es = = 555 tsf = 708 psi
5
ldquoDrained sandsiltrdquo gives a ν = 020
Step 4 Calculate the net cone tip resistance
qtnet = qt minus σvo = 1250 psi minus 189 psi = 12311 psi
Step 5 Calculate the bearing capacity of the sand
In drained sandssilts the bearing capacity is taken as the stress when (sB) = 011 (or 11 foundation width)
minus0345 minus0345 05 s L 600 qmax = hs ∙ qtnet ∙ ( ) ∙ ( ) = 078 ∙ (9795) ∙ (011)05 ∙ ( )
B B 144
= 193 psi = 27860 psf qmax
Assuming a factor of safety (FS) of 3
qmax = 645 psi = 9287 psf
3
Step 6 Calculate settlement
68
119902119898119886119909 0345 2
1 frasl119865119878 L 119904 = 119861 ∙ [ ∙ ∙ ( ) ]
ℎ119904 119902119905119899119890119905 B
2 0345 1 645 psi 600 inches s = 144 inches ∙ [ ∙ ∙ ( ) ] = 18 inches
078 12311 psi 144 inches
Step 7 Determine if q gt qmax
q = 8000 psf lt 9287psf = qmax3
69
CHAPTER 4 DIRECT CPT METHOD FOR DEEP FOUNDATIONS
41 INTRODUCTION
The axial compression capacity (Qtotal) for a single pile includes a side component (Qside) end bearing
component (Qbase) and pile weight (Wpile) as shown in Figure 30 Since piles commonly push through
several layers a summation of the unit side frictions acting on the pile segments must be considered
over the length of the pile While the examples shown in this Guide resemble hand calculations
computer software is more efficient Programming the procedure is possible and represents a practical
method of designing deep foundations However commercial software is frequently available and is
MnDOTrsquos most common method of designing deep foundations
There are upwards of 40 different direct CPT methods that have been developed over the past five
decades to determine a piles axial compression capacity Many of the earliest methods relied on hand-
recorded information where qc data from mechanical CPTs would be collected at 20 cm intervals
whereas the direct CPT method uses scaled penetrometer readings via specified algorithms to obtain
the pile unit side friction and unit end bearing The method that will be used for deep foundation design
is the Modified UniCone method which uses all three readings of the electronic piezocone (qt fs and u2)
while addressing a variety of pile foundation types
Figure 30 Direct CPT evaluation of axial pile capacity
70
The modified UniCone Method is based upon a total of 330 pile load tests (three times the original
Unicone database) that were associated with SCPTu data Originally the UniCone method provided
approximate soil classification in five groups via a chart of qE vs fs (Figure 31) where qE = qt -u2 Later
using the modified approach with a larger data set provided soil sub classifications as shown in Figure
32 This new 9-zone normalized soil behavior type is determined using CPT data in combination with the
CPT Material Index
Figure 31 UniCone Method soil behavior type using CPT (Mayne 2017)
Figure 32 Modified UniCone Method soil behavior type using CPT (Mayne 2017)
71
42 MODIFIED UNICONE METHOD
In order to determine the axial pile capacity using the modified method the first step requires the
determination of geoparameters as shown in Direct CPT Method for Soil Characterization Once the
soil unit weight and CPT Material index are determined for each soil layer then the effective cone
resistance pile unit side friction (fp) and pile end bearing resistance (qb) can be determined
421 Step 1
Work through the steps provided in CPT Method for Soil Characterization until each soil layer is
defined by its CPT material index and soil behavior type using Figure 6 After these steps have been
completed continue to Step 2 to determine qE qb and fp
422 Step 2
Once qt is determined for each layer the effective cone resistance can be calculated using Equation 37
119902119864 = 119902119905 minus 1199062 37
Where (a) qE is the specific value at each elevation along the pile sides for determining fp and (b) at the bottom of the pile qE is averaged in the vicinity of the pile tip from the tip bearing elevation to about one diameter beneath the tip for determining qb
Using CPT material index and qE the pile end bearing resistance is calculated using Equation 38
38 119902119887 = 119902119864 ∙ 10(0325∙119868119888minus1218)
The pile unit side friction is obtained from qE and Ic
39 119891119901 = 119902119864 ∙ 120579119875119879 ∙ 120579119879119862 ∙ 120579119877119860119879119864 ∙ 10(0732∙119868119888minus3605)
Where θPT = coefficient for pile type (084 for bored 102 for jacked 113 for driven piles) θTC = coefficient for loading direction (111 for compression and 085 for tension) θRATE = rate coefficient applied to soils in SBT zone 1 through 7 (109 for constant rate of penetration test and 097 for maintained load tests)
72
423 Step 3
Due to piles extending through multiple layers the unit side components acting on various pile
segments would need to be summed if not using the direct CPT method Since CPT calculates data at
regular intervals of 2 cm to 5 cm along the sides of the pile the average fp in each layer can be used
directly in Equation 40 to obtain the shaft capacity
119876119904119894119889119890 = 119891119901 ∙ 119860119904 40
Where fp = average pile side friction along pile length from eqn 39 As = πmiddotdmiddotH d = pile diameter and H = length embedded below grade
The base capacity for a pile in compression loading is given by Equation 41 For piles in tension (or uplift)
Qbase can be taken as 0
= 119902119887 ∙ 119860119887 41 119876119887119886119904119890 Where
qb = end bearing resistance from eqn 38 Ab = πmiddotd24 (area of a circular pile)
424 Step 4
The final step is to calculate the axial pile capacity using Equation 42
119876119905119900119905119886119897 = 119876119904119894119889119890 + 119876119887119886119904119890 minus 119882119901119894119897119890 42
43 AXIAL PILE DISPLACEMENTS
Movement of pile foundations can be assessed using elastic continuum theory which has been
developed using finite element analyses boundary elements and analytical closed-form solutions In
the case of piles passing through several soil layers the elastic solution can be used by stacking pile
segments (each with its own stiffness) as represented by soil Youngrsquos modulus The use of software is
recommended for pile groups Several available programs such as DEFPIG GROUP and PIGLET can
handle pile groups under axial and lateralmoment loading
73
44 EXAMPLE PROBLEMS
441 Example 5 Direct CPT Methods Axial Pile Capacity
Several piles need to be placed beneath the edge of a building Use the CPT data collected for this site
shown in Figure 34 to determine the axial capacity for one of the piles Assume round steel driven piles
will be used with a diameter of 1275 inches wall thickness of 025 inches and lengths of 80 feet
Concrete will be used to fill the piles To solve for all geoparameters use the Direct CPT Method for Soil
Characterization section All sand layers can be assumed ldquodrainedrdquo with ν = 02
Figure 33 Deep foundation end bearing pile diagram
74
Figure 34 CPT data from Minnesota for example problem 5
75
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 35 Soil layers using ldquorules of thumbrdquo for pile capacity example using direct CPT method
Layer 1
From Figure 35 fs = 18 psi taken as a representative value of the layer
76
18 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1210 pcf
145 psi
Layer 2
From Figure 35 fs = 12 psi taken as a representative value of the layer
12psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 3
From Figure 35 fs = 20 psi taken as a representative value of the layer
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
CPT Material Index
Layer 1
From Figure 35 fs = 18 psi and qc = 3000 psi
qt = qc + u2 ∙ (1 minus a) = 3000 psi + 20 psi ∙ (1 minus 08) = 3004 psi
∙ 4 feet = 1219 pcf ∙ 4 feet = 4877 psf = 34 psi σvo = γt
fs 18 psi Fr() = 100 ∙ = 100 ∙ = 060
(qt minus σvo) (3004 psi minus 34 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
77
uo = 0 psi
prime σvo = σvo minus uo = 34 psi minus 0 psi = 34 psi
(qt minus σvo)σatm (3004 psi minus 34 psi )145 psi = = Qtn prime )n (σvoσatm (34 psi145 psi)n
prime σvo 34 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(060)]2
Qtn = 3593 n = 038 lt 10 Ic = 14 (ie sand)
Layer 2
From Figure 35 fs = 12 psi and qc = 500 psi
qt = qc + u2 ∙ (1 minus a) = 500 psi + 40 psi ∙ (1 minus 08) = 508 psi
= 4838 psf + γt ∙ 45 feet = 4838 psf + 1172 pcf ∙ 45 feet = 5756 psf = 400 psi σvo
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 260
(qt minus σvo) (508 psi minus 400 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
78
uo = 0 psi
prime σvo = σvo minus uo = 400 psi minus 0 psi = 400 psi
(qt minus σvo)σatm (508 psi minus 400 psi )145 psi = = Qtn prime (σvoσatm)n (400 psi145 psi)n
prime σvo 400 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(260)]2
Qtn = 117 n = 10 le 10 Ic = 29 (ie clayey silt)
Layer 3
From Figure 35 fs = 20 psi and qc = 5000 psi
qt = qc + u2 ∙ (1 minus a) = 5000 psi + 0 psi ∙ (1 minus 08) = 5000 psi
= 5756 psf + γt ∙ 6 feet = 5756 psf + 1219 pcf ∙ 6 feet = 6488 psf = 451 psi σvo
fs 20 psi Fr() = 100 ∙ = 100 ∙ = 040
(qt minus σvo) (5000 psi minus 451 psi)
79
Step 2 Iterate to solve for Ic Steps are not shown for brevity
prime σvo = σvo minus uo = 451 psi minus 0 psi = 451 psi
(qt minus σvo)σatm (5000 psi minus 451 psi )145 psi = = Qtn prime )n (σvoσatm (451 psi145 psi)n
prime σvo 451 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(040)]2
= 1801 n = 06 lt 10 = 15 (ie sand) Qtn Ic
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic (14) Qtn (359) and Fr (06) the first layer is defined as a ldquoGravelly Sandrdquo
from Figure 36
80
Figure 36 Soil layer 1 using SBT method
81
Layer 2
Based on values of Ic (29) Qtn (117) and Fr (26) the second layer is defined as a ldquoSilty Mixrdquo
from Figure 37
Figure 37 Soil layer 2 using SBT method
82
Layer 3
Based on values of Ic (15) Qtn (180) and Fr (04) the third layer is defined as a ldquoSandrdquo from
Figure 38
Figure 38 Soil layer 3 using SBT method
Effective Cone Resistance
Layer 1
83
qE = qt minus u2 = 3004 psi minus 20 psi = 2984 psi
Layer 2
qE = qt minus u2 = 508 psi minus 40 psi = 468 psi
Layer 3
qE = qt minus u2 = 5000 psi minus 0 psi = 5000 psi
Pile End Bearing Resistance
Layer 3
= qE ∙ 10(0325∙Icminus1218) qb = 5000 psi ∙ 10(0325∙15minus1218) = 9085 psi
Pile Unit Side Friction
Layer 1
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 2984 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙14minus3605) = 99 psi
Layer 2
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 468 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙29minus3605) = 211 psi
Layer 3
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 5000 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙15minus3605) = 202 psi
84
Axial Pile Capacity
The side capacity is based on pile design as shown in Figure 39
Layer 1
Qside = fp ∙ As = 99 psi ∙ (π ∙ d ∙ L) = 99 psi ∙ (π ∙ 1275 in ∙ 48 in) = 19064 lb
Layer 2
Qside = fp ∙ As = 211 psi ∙ (π ∙ d ∙ L) = 211 psi ∙ (π ∙ 1275 in ∙ 588 in) = 457122 lb
Layer 3
Qside = fp ∙ As = 202 psi ∙ (π ∙ d ∙ L) = 202 psi ∙ (π ∙ 1275 in ∙ 240 in) = 58170 lb
Figure 39 Soil layering compared to pilings
85
End Bearing Resistance in Layer 3
d2 1275 in2
Qbase = qb ∙ Ab = 9085 psi ∙ (π ∙ ) = 9085 psi ∙ (π ∙ ) = 115932 lb 4 4
Wp = (γsteel ∙ Asteel ∙ Depth) + (γsteel ∙ Aconc ∙ Depth)
Wp = (490 pcf ∙ 003 ft2 ∙ 55 ft) + (150 pcf ∙ 085 ft2 ∙ 55 ft) = 7724 lbs
Layer 3
Qtotal = ΣQside + Qblayer3 minus Wp
Qtotal = (19064 + 45713 + 58170) lbs + 115932 lbs minus 7724 lbs = 642563 lbs
= 642 kips Qtotal
Table 8 Summary of selected and calculated parameters
Layer L (in) qc
(psi)
fs
(psi)
Fr Ic Qtn qE
(psi)
qb
(psi)
Qbase
(lb)
fp
(psi)
Qside
(lb)
Qtotal
(kip)
1 48 3000 18 06 14 356 2984 NA NA 99 19064
2 588 500 12 26 29 117 468 NA NA 211 457122
3 240 5000 20 04 15 180 5000 9085 115932 202 58170
Total 115932 534355 642
86
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Agaiby SA (2018) Advancements in the interpretation of seismic piezocone tests in clays and other geomaterials PhD dissertation School of Civil amp Environmental Engineering Georgia Institute of Technology Atlanta GA
Agaiby SS and Mayne PW (2018) Interpretation of piezocone penetration and dissipation tests in sensitive Leda Clay at Gloucester Test Site Canadian Geotechnical Journal (accepted for publication 08 March 2018) CGJ-2017-0388
Amar S Baguelin F Canepa Y and Frank R (1998) New design rules for the bearing capacity of shallow foundations based on Menard PMT Geotechnical Site Characterization Vol 2 (Proc ISC-1 Atlanta) 727-733
American Petroleum Institute (API) (1981) Planning designing and constructing fixed offshore platforms Recommended practice API-RP2A 17th edition Washington DC API
American Petroleum Institute (API) (1987) Planning designing and constructing fixed offshore platforms Recommended practice API-RP2A 18th edition Washington DC API
American Petroleum Institute (API) (1989) Planning designing and constructing fixed offshore platforms Recommended practice API-RP2A 19th edition Washington DC API
Andersen KH and Stenhamar P (1982) Static plate loading tests on overconsolidated clay Journal of the Geotechnical Engineering Division ASCE 108 (GT7) 919-934
Basu D and Salgado R (2012) Load and resistance factor design of drilled shafts in sands Journal of Geotechnical amp Geoenvironmental Engineering 138 (12) 1455-1469
Berardi R Jamiolkowski M and Lancellotta R (1991) Settlement of shallow foundations in sands selection of stiffness on the basis of penetration resistance Geotechnical Engineering Congress Vol 1 185-200 (Proc GSP-27 Boulder) Reston Virginia ASCE
Brand EW Muktabhant C and Taechathummarak A (1972) Load tests on small foundations in soft clay Performance of Earth and Earth-Supported Structures Vol 1 Part 2 903-928 (Proc PC) ASCE Reston Virginia ASCE
Briaud JL and Gibbens RM (1999) Behavior of five large spread footings in sand Journal of Geotechnical amp Geoenvironmental Engineering 125 (9) 787-797
Briaud JL (2007) Spread footings in sand load-settlement curve approach Journal of Geotechnical amp Geoenvironmental Engineering 133 (8) 905-920
Brown DA Turner JP and Castelli RJ (2010) Drilled Shafts Construction Procedures and LRFD Design Methods Geotechnical Engineering Circular GEC 10 Report No FHWA NHI-10-016 Washington DC National Highway Institute US DOT
87
Burland J B (1973) Shaft Friction of Piles in Clay -A Simple Fundamental Approach Ground Eng 6(3) 30-42
Cai G Liu S and Puppala A (2012) Reliability assessment of CPTu-based pile capacity predictions in soft clay deposits Engineering Geology 84-91
Canadian Geotechnical Society (1992) Canadian Foundation Engineering Manual Third Edition Vancouver British Columbia BiTech Publishers
Cerato AB and Lutenegger AJ (2007) Scale effects of shallow foundation bearing capacity on granular material Journal of Geotechnical amp Geoenvironmental Engineering 133 (10) 1192-1201
Chen BS-Y and Mayne PW (1996) Statistical relationships between piezocone measurements and stress history of clays Canadian Geotechnical Journal 33 (3) 488-498
Cho W and Finno RJ (2010_ Stress-strain responses of block samples of compressible Chicago glacial clays Journal of Geotechnical amp Geoenvironmental Engineering 136 (1) 178-188
Chung SF Randolph MF and Schneider JA (2006) Effect of penetration rate on penetrometer resistance in clay Journal of Geotechnical amp Geoenvironmental Engineering 132 (9) 1188-1196
Clausen CJF Aas PM and Karlsrud K (2005) Bearing capacity of driven piles in sand the NGI approach Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis Group
Consoli NC Schnaid F and Milititsky J (1998) Interpretation of plate load tests on residual soil site Journal of Geotecnical amp Geoenvironmental Engingeering 124 (9) 857-867
Dasenbrock DD (2006) Assessment of pile capacity by static and dynamic methods to reconcile design predictions with observed performance Proc 54th Minnesota Geotechnical Engineering Conference 95-108 edited by Labuz J Univ of Minnesota St Paul
DeBeer E and Martens A (1957) Method of computation of an upper limit for the influence of heterogeneity of sand layers on the settlement of bridges 275-282 Proc ICSMFE-4 Vol 1 London ICSMFE
DeBeer EE (1965) Bearing capacity and settlement of shallow foundations on sand Proc Symposium on Bearing Capacity and Settlement of Foundations 15-33 Duke University Durham NC
Decourt L (1999) Behavior of foundations under working load conditions Proc PCSMGE-11 Foz do Iguassu Brazil Vol 4 453-487
Demers D and Leroueil S (2002) Evaluation of preconsolidation pressure and the overconsolidation ratio from piezocone tests of clay deposits in Quebec Canadian Geotechnical Journal 39 (1) 174-192
Eslaamizaad S and Robertson PK (1996) Cone penetration test to evaluate bearing capacity of foundations in sand 429-438 Proc CGCFG-49 Vol 1 St Johns Newfoundland
Eslami A Alimirzaei M Aflaki E and Molaabasi H (2017) Deltaic soil behavior classification using CPTu records Marine Georesources amp Geotechnology 35 (1) 62-79
88
Eslami A and Gholami M (2005) Bearing capacity analysis of shallow foundations from CPT data 1463-1466 Proc ICSMGE- 16 Vol 3 (Osaka) Millpress Rotterdam
Eslami A and Gholami M (2006) Analytical model for the ultimate bearing capacity of foundations from cone resistance Scientia Iranica 13 (3) 223-233
Eslami A and Fellenius BH (1997) Pile capacity by direct CPT and CPTu methods applied to 102 case histories Canadian Geotechnical Journal 34 (6) 886-904
Fellenius BH (2009) Basics of Foundation Design Vancouver BiTech Publishers
Fellenius BH and Altaee A (1994) Stress and settlement of footings in sand Vertical and Horizontal Deformations of Foundations and Embankments 2 1760-1773
Fellenius BH and Eslami A (2000) Soil profile interpreted from CPTu data Proc Geotechnical Engineering Conference Year 2000 Geotechnics Asian Inst of Technology Bangkok Thailand Available from wwwfelleniusnet
Finno R J and Chung C K (1992) Stress-strain-strength responses of compressible Chicago glacial clays Journal of Geotechnical Engineering 118 (10) 1607ndash1625
Fleming K Weltman A Randolph M and Elson K (2009) Piling Engineering Third Edition London Taylor amp Francis Group
Frank R and Magnan J-P (1995) Cone penetration testing in France national report Proceedings International Symposium on Cone Penetration Testing 3 (CPT95 Linkoumlping) Swedish Geotechnical Society Report 3(95) 147-156
Gavin K Adekunte A and OKelly B (2009) A field investigation of vertical footing response on sand Geotechnical Engineering 162 257-267
Hegazy YA (1998) Delineating geostratigraphy by cluster analysis of piezocone data PhD dissertation School of Civil amp Environmental Engineering Georgia Institute of Technology Atlanta GA
Holtz RD Kovacs WD and Sheehan TC (2011) An Introduction to Geotechnical Engineering 2nd Edition London Pearson Publishing
Hunt CE Pestana JM Bray JD and Riemer M (2002) Effect of pile driving on static and dynamic properties of soft clay Journal of Geotechnical amp Geoenvironmental Engineering 128 (1) 13-24
Jamiolkowski M Ladd CC Germaine JT and Lancellotta R (1985) New developments in field and lab testing of soils Proc ICSMFE-11 1 57-154
Jardine R Chow F Overy R and Standing J (2005) ICP design methods for driven piles in sands and clays London Thomas Telford Publishing
Jardine RJ Lehane BM Smith P and Gildea P (1995) Vertical loading experiments on rigid pad foundations at Bothkennar Geotechnique 45 (4) 573-597
89
Karlsrud K (2012) Prediction of load-displacement behavior and capacity of axially-loaded piles in clay based on analyses and interpretation of pile load test results PhD Dissertation Norwegian Univ Science amp Technology Trondheim Norway
Karlsrud K Clausen CJF and Aas PM (2005) Bearing capacity of driven piles in clay the NGI approach Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis
Karlsrud K Lunne T Kort DA and Strandvik S (2001) CPTU correlations for clays Proc 16th ICSMGE 2 683-702
Kimura T Kusakabe O and Saitoh K (1985) Geotechnical model tests of bearing capacity problems in a centrifuge Geotechnique 35 (1) 33-45
Knappett JA and Craig RF (2012) Craigs Soil Mechanics 8th Edition London Spon Press Taylor amp Francis
Kolk HJ and Velde EVD (1996) A reliable method to determine friction capacity of piles driven into clays Paper No 7993 Proc Offshore Technology Conference Houston
Kolk HJ Baaijens AE and Senders M (2005) Design criteria for pipe piles in silica sands Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis Group
Kulhawy FH (2004) On the axial behavior of drilled shafts Geosupport 2004 1-18
Kulhawy FH and Jackson CS (1989) Foundation Engineering Current Principles amp Practices 2 (GSP 22) Reston VA ASCE
Kulhawy FH and Mayne PW (1990) Manual for Estimating Soil Properties for Foundation Design EPRI Report EL-6800 Electric Power Research Institute Palo Alto 306 page Download from wwwepricom
Kulhawy FH Trautmann CH Beech JF ORourke TD and McGuire W (1983) Transmission Line Structure Foundations for Uplift-Compression Loading Report EL-2870 Palo Alto CA Electric Power Research Institute wwwepricom
Ku T and Mayne PW (2015) In-situ lateral stress coefficient (K0) from shear wave velocity measurements in soils Journal of Geotechnical amp Geoenvironmental Engineering 141 (12) 101061(ASCE) GT1943-56060001354
Ladd CC and DeGroot DJ (2003) Recommended practice for soft ground site characterization Soil amp Rock America 2003 3-57
Larsson R (1997) Investigations and Load Tests in Silty Soils SGI Report R-54 Linkoumlping Swedish Geotechnical Institute
Larsson R (2001) Investigations and Load Tests in Clay Till SGI Report R-59 Linkoumlping Swedish Geotechnical Institute
Lee J and Salgado R (2005) Estimation of bearing capacity of circular footings on sands based on cone penetration tests Journal of Geotechnical amp Geoenvironmental Engineering 131 (4) 442-452
90
Lehane BM (2003) Vertically loaded shallow foundation on soft clayey silt Geotechnical Engineering 156 (1) Thomas Telford London 17-26
Lehane BM (2013) Relating foundation capacity in sands to CPT qc Geotechnical amp Geophysical Site Characterization 1 London Taylor amp Francis Group
Lehane BM Doherty JP and Schneider JA (2008) Settlement prediction for footings on sand Deformational Characteristics of Geomaterials (IS-Atlanta) 1 Rotterdam Millpress
Lehane BM Li Y and Williams R (2013) Shaft capacity of displacement piles in clay using the CPT Journal of Geotechnical amp Geoenvironmental Engineering 139 (2) 253-266
Lehane BM Schneider JA and Xu X (2005) The UWA-05 method for prediction of axial capacity of driven piles in sand Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis
Low HE Lunne T Andersen KH Sjursen MA Li X and Randolph MF 2010 Estimation of intact and remoulded undrained shear strengths from penetration tests in soft clays Geotechnique 60 (11) 843-859
Lunne T Robertson PK and Powell JJM (1997) Cone Penetration Testing in Geotechnical Practice London RoutledgeBlackie Academic
Lutenegger AJ and DeGroot DJ (1995) Settlement of Shallow Foundations on Granular Soils Report Contract 6332 Task Order 4 submitted to Massachusetts Highway Dept by University of Mass-Amherst Amherst MA
Lutenegger AJ and Adams MT (2003) Characteristic load-settlement behavior of shallow foundations Proc FONDSUP 2 381-392
Mase T and Hashiguchi K (2009) Numerical analysis of footing settlement problem by subloading surface model Soils amp Foundations 49 (2) 207-220
Mayne PW (1987) Determining preconsolidation stress and penetration pore pressures from DMT contact pressures Geotechnical Testing Journal 10(3) 146-150
Mayne PW (1991) Determination of OCR in Clays by Piezocone Tests Using Cavity Expansion and Critical State Concepts Soils and Foundations 31 (2) 65-76
Mayne PW (2007) NCHRP Synthesis 368 on Cone Penetration Test Washington DC Transportation Research Board National Academies Press wwwtrborg
Mayne PW (2009) Geoengineering Design Using the Cone Penetration Test Richmond BC Canada ConeTec Inc
Mayne PW (2014) Keynote Interpretation of geotechnical parameters from seismic piezocone tests Proceedings 3rd Intl Symp on Cone Penetration Testing 47-73
Mayne PW (2016) Evaluating effective stress parameters and undrained shear strengths of soft-firm clays from CPT and DMT Australian Geomechanics Journal 51 (4) 27-55
91
Mayne PW (2017) Stress history of soils from cone penetration tests 34th Manual Rocha Lecture Soils amp Rocks Vol 40 (3) Saouml Paulo ABMS
Mayne PW Christopher B Berg R and DeJong J (2002) Subsurface Investigations -Geotechnical Site Characterization Publication No FHWA-NHI-01-031 Washington DC National Highway Institute Federal Highway Administration
Mayne PW Coop MR Springman S Huang A-B and Zornberg J (2009) Geomaterial Behavior and Testing Proc 17th Intl Conf Soil Mechanics amp Geotechnical Engineering 4 Rotterdam MillpressIOS Press
Mayne PW and Illingworth F (2010) Direct CPT method for footings on sand using a database approach Proc ISCPT-2 3 315-322
Mayne PW and Niazi F (2017) Recent developments and applications in field investigations for deep foundations Proceedings 3rd Bolivian Conference on Deep Foundations 1 141-160
Mayne PW Peuchen J and Baltoukas D (2015) Piezocone evaluation of undrained strength in soft to firm offshore clays Frontiers in Offshore Geotechnics III 2 1091-1096
Mayne PW and Poulos HG (1999) Approximate displacement influence factors for elastic shallow foundation systems ASCE Journal of Geotechnical amp Geoenvironmental Engineering 125 (6) 453-460
Mayne PW Uzielli M and Illingworth F (2012) Shallow footing response on sands using a direct method based on cone penetration tests Full Scale Testing and Foundation Design (GSP 227) Reston Virginia ASCE
Mayne PW and Woeller DJ (2014) Direct CPT method for footings on sands silts and clays Geocharacterization for Modeling and Sustainability (GSP 234) 1983-1997
McClelland B (1974) Design of Deep Penetration Piles for Ocean Structures Journal Geot Eng Div 100(GT7) 705-747
Mesri G (1975) Discussion on new design procedure for stability of soft clays ASCE Journal of the Geotechnical Engineering Division 101(GT4) pp 409-412
Meyerhof GG (1956) Penetration tests and bearing capacity of cohesionless soils Journal of Soil Mechanics amp Foundations Division 82 (SM1) 1-19
Meyerhof GG (1974) General Report Penetration testing outside Europe Proceedings ESOPT-1 Vol 21 Swedish Geotechnical Society Stockholm
Meyerhof GG (1976) Bearing capacity and settlement of pile foundations Journal of Geotechnical Engineering Division 102(3) 197-228
Missiaen T Verhegge J Heirman K and Crobe P (2015) Potential of cone penetrating testing for mapping deeply buried palaeolandscapes in the context of archaeological surveys in polder areas Journal of Archaeological Science 55 174-187
92
Mlynarek Z Wierzbicki J Goglik S and Bogucki M (2014) Shear strength and deformation parameters of peat and gyttja from CPTu DMT and VST Proceedings 5th International Workshop on CPTu and DMT in soft clays and organic soils 193-210 Poznan Poland Polish Committee on Geotechnics Menard Polska Tensar Internationa
Nejaim PF Jannuzzi GMF and Danziger FAB (2016) Soil behavior type of the Sarapuiacute II test site Geotechnical amp Geophysical Site Characterization 5 1009-1014
Niazi FS and Mayne PW (2013) Cone penetration test based direct methods for evaluating static axial capacity of single piles Geotechnical and Geological Engineering 31 (4) 979-1009
Niazi FS and Mayne PW (2015) Enhanced Unicone expressions for axial pile capacity evaluation from piezocone tests Proc International Foundations Conference amp Equipment Expo RestonVA ASCE
Niazi FS and Mayne PW (2016) CPTu-based enhanced UniCone method for pile capacity Engineering Geology 212 21-34
Nowacki F Karlsrud K and Sparrevik P (1996) Comparison of recent tests on OC clay and implications for design Large Scale Pile Tests in Clay London Thomas Telford
ONeill MW (2001) Side resistance in piles and drilled shafts Journal of Geotechnical amp Geoenvironmental Engineering 127 (1) 3-16
Ouyang Z and Mayne PW (2018) Effective friction angle of clays and silts from piezocone Canadian Geotechnical Journal in press doiorg101139cgj-2017-0451
Paikowsky SG Canniff MC Lesny K Kisse A Amatya S and Muganga R (2010) LRFD Design and Construction of Shallow Foundations for Highway Bridge Structures NCHRP Report 651 Washington DC National Cooperative Highway Research Program Transportation Research Board
Pestana JM Hunt CE and Bray JD (2002) Soil deformation and excess pore pressure filed around a closed-ended pile Journal of Geotechnical amp Geoenvironmental Engineering 128 (1) 1-12
Poulos HG and Davis EH (1974) Elastic Solutions for Soil and Rock Mechanics New York Wiley amp Sons
Poulos HG and Davis E (1980) Pile Foundation Analysis amp Design New York John Wiley amp Sons
Powell JJM and Lunne T (2005) Use of CPTU data in clays and fine-grained soils Studia Geotechnica et Mechanica XXVII(3) 29-66
Randolph MF (2003) Rankine Lecture Science and empiricism in pile foundation design Geotechnique 53 (10) 847-875
Robertson PK (1990 1991) Soil classification using the cone penetration test Canadian Geotechnical Journal 27 (1) 151-158 Closure 28 (1) 176-178
Robertson PK (1991) Estimation of foundation settlements in sand from CPT Geotechnical Engineering Congress 2 Reston Virginia ASCE
93
Robertson PK (2009) Cone penetration testing a unified approach Canadian Geotechnical J 46 (11) 1337-1355
Robertson PK (2009a) Performance based earthquake design using the CPT Keynote Lecture International Conference on Performance-Based Design in Earthquake Geotechnical Engineering IS Tokyo Tsukuba Japan
Robertson PK (2009b) Interpretation of cone penetration tests a unified approach Canadian Geotechnical Journal 46 (11) 1337-1355
Robertson PK and Cabal KL (2007) Guide to cone penetration testing Second Edition Signal Hill California Gregg In-Situ
Robertson PK and Cabal KL (2015) Guide to Cone Penetration Testing for Geotechnical Engineering 6th Edition Signal Hill California Gregg Drilling amp Testing Inc
Schmertmann JH (1970) Static cone to compute static settlement over sand Journal of the Soil Mechanics and Foundations Division 95 (SM3) 1011-1043
Schmertmann JH (1978) Guidelines for cone penetration test performance and design Report FHWA-TS-78-209 Washington DC Federal Highway Administration
Schnaid F Wood WR Smith AKC and Jubb P (1993) Investigation of bearing capacity and settlements of soft clay deposits at Shellhaven Predictive Soil Mechanics London Thomas Telford
Schneider JA Xu X and Lehane BM (2008) Database assessment of CPT-based design methods for axial capacity of driven piles in siliceous sands J Geotechnical and Geoenvironmental Engineering 134 (9) 1227-1244
Semple D (1980) Recent Developments in the Design amp Construction of Piles London Institution of Civil Engineers
Senneset K Sandven R amp Janbu N (1989) Evaluation of soil parameters from piezocone tests In-Situ Testing of Soil Properties for Transportation Transportation Research Record 1235 24-37
Shahri AA Melehmir A and Juhlin C (2015) Soil classification analysis based on piezocone penetration test data Engineering Geology 189 32-47
Stuedlein AW and Holtz RD (2010) Undrained displacement behavior of spread footings in clay The Art of Foundation Engineering Practice GSP 198 Reston Virginia ASCE
Takesue K Sasao H and Matsumoto T (1998) Correlation between ultimate pile skin friction and CPT data Geotechnical Site Characterization 2 1177ndash1182
Tand KE Funegard EG and Briaud J-L (1986) Bearing capacity of footings on clay CPT method Use of In-Situ Tests in Geotechnical Engineering GSP 6 1017-1033
Tand KE Warden PE and Funegaringrd EG (1995) Predicted-measured bearing capacity of shallow footings on sand PISCPT 2 589-594
Thomas D (1968) Deep soundings test results and the settlement of spread footings on normally consolidated sands Geotechnique 18 (2) 472-488
94
Tomlinson MJ (1957) The adhesion of piles driven into clay soils Proc 4th Intl Conf on Soil Mechanics and Foundation Engineering 2 66-71
Tomlinson MJ (1986) Foundation Design amp Construction London Longman Scientific
Tomlinson MJ (1987) Pile Design and Construction Practice London Viewpoint Publication
Trofimenkov JG (1974) Penetration testing in the USSR Proceedings ESOPT-1 1 147-154
Tumay MT Hatipkarasulu Y Marx ER and Cotton B (2013) CPTPCPT-based organic material profiling Proc 18th Intl Conf on Soil Mechanics amp Geotechnical Engineering Paris 633-636
Uzielli M and Mayne PW (2011) Statistical characterization and stochastic simulation of load-displacement behavior of shallow footings Proc Georisk 2011 Geotechnical Risk Management amp Assessment (GSP 224) 672-679
Uzielli M and Mayne PW (2012) Load-displacement uncertainty of vertically-loaded shallow footings on sands and effects on probabilistic settlement estimation Georisk Assessment and Management of Risk for Engineered Systems and GeoHazards 6 (1) 50-69
Valsson SM (2016) Detecting quick clay with CPTu Proceedings 17th Nordic Geotechnical Meeting Reykajavik Iceland Challenges in Nordic Geotechnics
Van Dijk BFJ and Kolk HJ (2011) CPT-based design method for axial capacity of offshore piles in clay Frontiers in Offshore Geotechnics II 555-560
Vesić AS (1975) Bearing capacity of shallow foundations Foundation Engineering Handbook New York Van Nostrand Reinhold Publishers
Vesic AS (1977) NCHRP Synthesis of Highway Practice 42 Design of Pile Foundations Washington DC Transportation Research Board
Yu F and Yang J (2012) Base capacity of open-ended steel pipe piles in sand J Geotechnical and Geoenvironmental Engineering 138 (9) 1116-1128
Zawrzykraj P Rydelek P and Bakowska A (2017) Geoengineering properties of Eemian peats from Radsymin (central Poland) in the light of static cone penetration and dilatometer tests Engineering Geology 226 290-300
95
APPENDIX A
List of Figures
Figure A1 Conventional methods versus direct CPT approach to shallow foundation response A-3
Figure A2 Direct bearing capacity relationship for embedded square and strip footings situated on clean
sands (after Schmertmann 1978) A-6
Figure A3 Direct CPT bearing factor Rk as function of footing width B and embedment depth from finite
element analyses (Tand et al 1995) A-6
Figure A4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio and sand
consistency (Eslaamizaad amp Robertson 1996) A-7
Figure A5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp Salgado (2005) in
terms of footing size relative density and base settlement A-8
Figure A6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and normalized
cone tip resistance (after Eslami amp Gholami 2005) A-8
Figure A7 Direct relationship between ultimate bearing stress in clays and measured cone tip resistance
(Schmertmann 1978) A-10
Figure A8 Direct relationship between ultimate foundation bearing stress in clays and net cone tip
resistance (after Tand et al 1986) A-11
Figure A9 Measured load-displacements for each of the five large footings at TAMU (Briaud amp Gibbens
1994) sand site A-12
Figure A10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU (Briaud amp
Gibbens 1994) footings A-13
Figure A11 Characteristic stress versus square root normalized displacements for TAMU footings A-15
Figure A12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sand A-16
Figure A13 Summary of sand formation factors with corresponding CPT resistances A-19
Figure A14 Normalized foundation stresses vs square root of normalized displacement for spread
footings on sand (modified after Mayne et al 2012) A-20
Figure A15 Definitions of capacity from three types of observed foundation load test responses
(modified after Kulhawy 2004) A-21
Figure A16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footings A-22
A-1
Figure A17 Measured load vs displacement curves for 3 shallow foundations on clay at Baytown Texas
(Stuedlein amp Holtz 2010) A-25
Figure A18 Characteristic stress vs square root of normalized displacement for all three foundations at
Baytown site (Stuedlein amp Holtz 2010) A-26
Figure A19 Normalized foundation stress vs square root of normalized displacements A-27
Figure A20 Applied foundation stress vs CPT-calculated stress for 67 footings A-28
Figure A21 Applied footing stress vs CPT calculated stress on logarithmic scales A-29
Figure A22 Summary graph and equations of direct CPT method for evaluating the vertical A-30
Figure A23 Foundation soil formation parameter hs versus CPT material index Ic A-32
Figure A24 Bearing capacity ratio qmaxqnet versus CPT material index Ic A-33
Figure A25 Influence factor for rectangular foundations from elastic theory solution (Mayne amp
Dasenbrock 2017) A-34
List of Tables
Table A1 Direct CPT methods for bearing capacity of footings on clean sands A-4
Table A2 Direct CPT methods for bearing capacity of footings on clays A-9
Table A3 Summary of large footings soil conditions and reference sources of database A-17
Table A4 Sources for shallow foundation performace A-34
A-2
A1 INTERPRETATION OF SOIL PARAMETERS FROM CONE
PENETRATION TESTS
A11 Introduction
Geotechnical site characterization is an important first step towards the evaluation of
subsurface conditions and determination of soil layering geomaterial classification and the
evaluation of soil engineering parameters for the analysis and design of foundations retaining
walls tunnels excavations embankments and slope stability Increasing use of the electronic
cone penetrometer in highway applications is prominent in the USA because the results are
obtained much faster and less expensive than traditional methods that rely on rotary drilling
augering sampling and laboratory testing Moreover the cone penetration test (CPT) provides
at least three independent and continuous readings on soil behavior that are digitally recorded
and fully available at the end of the sounding As such the reliable and consistent
interpretation of CPT data is important since civil engineering works will best realize the
efficiency and economy of this technology in constructed facilities for county state and
interstate projects
A111 Cone Penetration Testing
The cone penetrometer is an electronically-instrumented steel probe that is vertically pushed into the
ground at constant rate of 20 mms (08 insec) using a hydraulic system The penetrometer is outfitted
with load cells pressure transducers and inclinometers that measure at least three readings
continuously with depth (a) cone tip resistance qt (b) sleeve friction fs and (c) porewater pressure u2
With the latter reading the device is called a piezocone thus the designation CPTu is used to indicate
porewater pressure Additional sensors are available that can be used to measure inclination
resistivity pH temperature shear wave velocity and other readings if desired
Figure A1 shows a schematic rendering of the cone penetration test that is performed in the
field per ASTM D 5778 procedures The hydraulic system is nominally rated at 180 kN (20 tons)
capacity and often mounted on a truck but can also be positioned on tracked vehicles or
anchored frames Figure A2 shows a MnDOT cone truck that uses the full dead weight of the
rig for the hydraulic reaction forces which are applied at the center of the vehicle The cab is
enclosed so that soundings may proceed during inclement weather and the data acquisition
system and operator are protected
Figure A 1 Setup and procedures for co ne penetration testing (CPT)
A-4
Figure A2 CPT truck-mounted rig used by MnDOT
A representative sounding taken at the I-35 test site just northeast of Saint Paul is presented in Figure
A3 Here the separate profiles of qt fs and u2 with depth are shown in side-by-side plots In the last
graph the hydrostatic porewater pressure (uo) due to the groundwater table is indicated by the blue
dashed line
These readings are used to interpret the soil layers types and properties of the ground as
discussed in the following sections
A-5
Figure A3 Representative CPT sounding at I-35 test site northeast of St Paul MN showing profiles of
(a) cone resistance (b) sleeve friction and (c) penetration porewater pressures
A112 Soil Parameter Interpretation
A variety of soil engineering parameters can be interpreted from CPT results on the basis of
theoretical frameworks analytical models or numerical simulations otherwise by empirical
methods based on correlations and statistics of databases Figure A4 shows a conceptual
utilization of the readings from the CPT for interpretation of geostratigraphy compressibility
flow characteristics and stress-strain-strength behavior of soils
Table A1 lists a selection of geoparameters that have been addressed for these purposes The
most common or useful relationships will be discussed in subsequent sections and reference is
made to other sources (Lunne et al 1997 Mayne 2007 Robertson amp Cabal 2015) for additional
details
A-6
Figure A4 Conceptual post-processing of CPT results for geoparameter evaluation
Table A1 List of common geoparameters interpreted from CPT data
Symbol Parameter Remarks Notes
SBT Soil behavior type (SBT)
Ic CPT material index
γt Unit weight
ρt Mass Density
σvo Overburden stress
u0 Hydrostatic pressure
A-7
σvo Effective stress
Table A1 Continued
Symbol Parameter Remarks Notes
σp Preconsolidation stress σp = 033 (qnet)m where m = fctn (Ic)
(or Effective Yield Stress) where stresses are in kPa
YSR Yield stress ratio YSR = σpσvo (formerly OCR)
c Effective cohesion intercept
ϕ Effective friction angle
su Undrained shear strength
D Constrained modulus
G0 Small-strain modulus
G Shear modulus
E Youngs modulus
cvh Coefficient of consolidation
K Bulk modulus
A-8
LI
K0 Lateral stress coefficient
k Hydraulic conductivity
Liquefaction index
A12 Geostratigraphy and Soil Behavioral Type (SBT)
In routine CPT soil samples are not normally taken and thus the measured stress friction andor
porewater pressure readings are used to infer the types of soils This can be accomplished using
three basic approaches
a Rules of Thumb or approximate guidelines for a quick visual assessment
b Soil Behavioral Type (SBT) Charts that are based on either the three readings (qt fs u2) net readings
including net cone resistance (qnet = (qt-σvo) effective cone resistance (qE = qt - u2) friction ratio (Rf =
100middotfsqt) and excess porewater pressures or using normalized CPT readings such as Q F Bq or U
c Probabilistic methods where the CPT readings have been calibrated from lab tests on
recovered soil samples (eg Tumay et al 2013)
A121 Approximate Rules of Thumb
The approximate rules of thumb provide simple guidelines for soil type and suggest that sands are
identified when qt gt 5 MPa and u2 asymp u0 whereas intact clays are prevalent when qt lt 5 MPa and u2 gt u0
(Mayne et al 2002) The magnitude of porewater pressures reflects the consistency of the intact clay
such that soft (u2 asymp 2middotu0) firm (u2 asymp 4middotu0) stiff (u2 asymp 8middotu0) and hard (u2 asymp 20middotu0) However for fissured
overconsolidated clays the measured porewater pressures are less than hydrostatic in fact often
negative u2 lt 0 On land negative porewater pressure readings at the shoulder filter element of the
piezocone should be greater than -1 bar ie u2 gt -100 kPa (-147 psi) Also for clean sands the friction
ratio (FR = 100middotfsqt lt 1) while for insensitive clays (FR gt 4)
A-9
Figure A5 presents a 36-m deep piezocone sounding from the MnDOT Wakota Bridge site which crosses
the Mississippi River at I-494 (Dasenbrock 2006) The qt and u2 readings have been annotated using the
aforementioned rules of thumb indicating the predominance of sands at this site As noted there are
five interbedded clay layers evident at depths of 1 m 3-4 m 7 - 12 m 17 m and 22-27 m
Figure A5 Interpretation of soil types from CPT data taken from the Wakota Bridge on I-494 using
approximate rules of thumb
A122 Soil Behavioral Type Charts
The most common approach for identification of soil types is based on soil behavioral charts
Reviews of several chart methods are provided elsewhere (Kulhawy amp Mayne 1990 Fellenius amp
Eslami 2000 Shahri et al 2015) Current popularity favors the 9-zone classification system to
evaluate soil behavioral type (SBT) that uses normalized piezocone readings (Robertson 1990
1991 Lunne et al 1997)
A-10
(119902119905 minus 120590119907119900) (a) normalized tip resistance 119876 = A11 prime 120590119907119900
100∙119891119904 (b) normalized sleeve friction 119865() = A12
(119902119905 minus 120590119907119900)
(1199062 minus 1199060) (c) normalized porewater pressure 119861119902 = A13
(119902119905 minus 120590119907119900)
While the Q-F-Bq classification is a three-dimensional plot it is often presented as a set of two-dimensional
plots specifically Q vs F and Q vs Bq as shown in Figure A6 Data are grouped according to layers and
can be superimposed to identify their association with the nine soil behavioral types All indirect CPT soil
classification approaches should be cross-checked and verified for a particular geologic setting and local
geotechnical conditions before routine use in practice
Figure A6 Nine-zone soil behavioral type charts (a) normalized cone resistance vs normalized
sleeve friction (b) normalized cone resistance vs normalized porewater pressure parameters
(adapted after Robertson 1991 Lunne et al 1997)
The development of a CPT material index (Ic) has been found advantageous in the initial screening of soil
types and helps to organize the sounding into their respective SBT zones of similar soil response In this
case the CPT index is found from (Robertson 2009)
A-11
A2 22 )log221()log473( rtnc FQI
where Qtn = modfied stress-normalized net cone resistance given by
(119902119905 minus 120590119907119900)120590119886119905119898 119876 = A31 prime (120590119907119900120590119886119905119898)119899
where σatm = reference stress equal to atmospheric pressure (1 atm asymp 1 bar = 100 kPa asymp 1 tsf asymp 147 psi)
The exponent n is soil-type dependent n = 1 (clays) n asymp 075 (silts) and n asymp 05 (sands) If units of bars are
used then Qtn (units of bars) is simply
(119902119905 minus 120590119907119900) 119876 = 119899 A32
prime ) (120590119907119900
The operational value of exponent n is found by iteration Initially an exponent n = 1 is used to calculate
the starting value of Ic (ie Qtn = Q) and then the exponent is upgraded to
prime 120590119907119900 119899 = 0381 ∙ 119868119888 + 005 ( ) minus 015 le 10 A4 120590119886119905119898
Then the index Ic is recalculated Iteration converges quickly and generally only 3 cycles are needed to
secure the operational Ic at each depth The soil zones and associated Ic values are detailed in Figure A7
The sensitive soils of zone 1 can be screened using the following expression
A-12
Zone 1 119876119905119899 lt 12119890119909119901( minus14 ∙ 119865119903)
A5
If any soil layers are found within zone 1 (sensitive soils) then caution should be exercised as these
structured clays are prone to instability collapses and difficulties in construction and performance
Another indicator of zone 1 sensitive soils is when the normalized porewater parameter Bq gt 08 When
these criteria are evident the geoengineer should consult with senior geotechnical personnel or the chief
engineer for guidance
Figure A7 Delineation of soil behavioral type zones using CPT material index Ic
A-13
Stiff overconsolidated clayey sands and sandy clays soils of zone 8 (15 lt Fr lt 45) and zone 9 (Fr gt
45) can be identified from the following criterion
Zones 8 and 9 0020)1(00030)1(0050
12
rr
tnFF
Q A6
The remaining soil types are identified by the CPT material index Zone 2 (organic clayey soils Ic ge 360)
Zone 3 (clays to silty clays 295 le Ic lt 360) Zone 4 (silt mixtures 260 le Ic lt 295) Zone 5 (sand mixtures
205 le Ic lt 260) Zone 6 (clean sands 131 le Ic lt 205) and Zone 7 (gravelly to dense sands Ic le 131)
The red dashed line at Ic = 260 represents an approximate boundary separating drained (Ic lt 260) from
undrained behavior (Ic gt 260)
Once the specific zone has been assigned to a layer a visual representation can be made to show
either the zone number or colorization so that the predominant layers and soil types can be
realized
A123 Probabilistic Soil Types from CPT
The inference of soil type has been addressed using probabilistic methods as discussed by Abu-
Farshakh et al (2008) and Tumay et al (2013) Here the CPT readings can be post-processed to
provide the percentage of sand silt and clay components with depth The output is consistent
and compatible with the current MnDOT soil classification chart (Figure A8) which is similar in
content to that used by the USDA
A-14
Figure A8 Chart for soil classification system by MnDOT
A13 Identifying Soft Organic Soils
The identification of soil type using cone penetration tests (CPT) is usually done on the basis of empirical
soil behavioral type (SBT) charts that use the measured readings (qt fs andor u2) While there at least
25 different sets of charts available (ie Hegazy 1998 Eslami et al 2017 Valsson 2016) one of the most
widely used and popular is the 9-zone SBT system that uses three normalized cone readings Qtn Fr and
Bq (Robertson amp Cabal 2007 Lunne et al 1997) The system is denoted as SBTn and plotted on charts in
terms of Q vs F and Q vs Bq as shown in Figure A9
A-15
1
10
100
1000
01 1 10
Norm
aliz
ed
Con
e T
ip R
esis
tan
ce
Q
Normalized Friction Fr = 100 fst(qt - svo) ()
9-Zone Soil Behavioral Type Chart
Gravelly Sands(zone 7)
Sands
(zone 6)
Sandy Mixtures(zone 5)
Silt Mix(zone 4)
Clays
(zone 3)
Organic Soils(zone 2)
Sensitive Clays
and Silts(zone 1)
Very stiff OC clayto silt (zone 9)
Very stiffOC sandto clayey
sand(zone 8)
1
10
100
1000
-06 -04 -02 00 02 04 06 08 10 12 14
No
rmal
ized
Co
ne
Tip
Res
ista
nce
Q
Normalized Porewater Bq = Du2(qt-svo)
Clays(zone 3)
Sensitive(zone 1)Organic
(zone 2)
Gravelly Sands(zone 7)
Sands(zone 6)
Figure A9 CPT charts for SBTn system (a) Q versus F (b) Q versus Bq
Of particular concern in geotechnical site characterization is the proper identification of soft
organic soils such as organic clays and silts (OL and OH) and peats (Pt) These soils often cause
difficulties and problems during construction and performance of civil engineering works because
of bio-degradation long-term creep excessive settlements andor other issues
The SBTn system has a category labelled Zone 2 recognizes organic soils However several recent
studies have noted that data from CPTu soundings do not always fall within the bounds
delineated using Zone 2 even though laboratory tests and field inspections clearly note the
presence of organic soils Lab tests to classify organic soils includes loss on ignition (LOI gt 5)
and two pairs of Atterberg limits testings per ASTM standards as well as the visual-manual
identification of strong organic odor and smell and dark color notably black and dark gray
Results by Mlynarek et al (2014) show that organic soils in Poland are not adequately identified by the
Q-F chart as seen in Figure A10 Moreover studies by Zawrzykraj et al (2017) found that neither the Q-
A-16
F and Q-Bq plots worked well to recognize organic soils and peats in Poland Nejaim et al (2016) found
that Brazilian soft organic clays were misclassified as Zone 3 (clays) rather than Zone 2 (organic clays)
Similarly a recent study on CPTu data from organic clays located at 24 sites in the USA Sweden Mexico
Brazil and Australia have been compiled (Agaiby 2018) These data are shown to generally avoid falling
with the Zone 2 boundaries in either Q-F or Q-Bq charts thus are not recognized during these soundings
as indicated by Figure A11 The exception in this case is the Mexico City clay which is correctly identified
however the other 23 sites are generally not properly recognized and categorized Additional findings by
Missiaen et al (2015) found that the CPTu results in Belgian soils also did not identify organic clays and
peats in the non-normalized version of these charts that uses Qt versus friction ratio (Rf = 100middotfsqt) as
detailed by Robertson amp Cabal (2007)
Figure A10 Soil behavioral chart with superimposed CPTu data from Polish organic soils
(from Mlynarek et al 2014)
A-17
Figure A11 Paired SBTn charts with CPTu data from 24 organic clay sites indicating non-compliance
with Zone 2 (organic soils)
(compiled by Agaiby 2018)
A14 Simplified Yield Stress Evaluation in Clays by CPTu
From the development of a hybrid formulation of spherical cavity expansion (SCE) theory with critical-
state soil mechanics (CSSM) the effective preconsolidation stress or yield stress (σp) of clays can be
expressed in terms of three piezocone parameters (a) net cone tip resistance qnet = qt - σvo (b)
measured excess porewater pressure Δu2 = u2 - u0 and (c) effective cone resistance qE = qt - u2 Details
on the SCE-CSSM solutions are given elsewhere (Mayne 1991 Mayne 2017 Agaiby amp Mayne 2018)
A-18
For normal and well-behaved clays which include inorganic fine-grained soils such as clays and silts
of low sensitivity a set of simple relationships can be derived By adopting characteristic default values
for effective friction angle ϕ = 30deg and rigidity index (IR = 100) linear expressions for the effective
preconsolidation stress are obtained
prime 120590119901 = A7 033 ∙ 119902119899119890119905
prime 120590119901 = 053 ∙ ∆1199062 A8
prime 120590119901 = 060 ∙ 119902119864 A9
A141 Application to soft Chicago clay
An example of a common geomaterial in this category includes the infamous soft Chicago clays
that were deposited in a glacio-lacustrine environment (Cho amp Finno 2010) The soft clay has a
mean water content of 20 liquid limit of 38 and plasticity index PI= 12 Results of CPTu tests
conducted at the national geotechnical experimentation site (NGES) at Northwestern University
are shown in Figure A12 (Ouyang amp Mayne 2017) As seen in the profile a thin soft clay resides
between depths of 9 to 105 m with a thicker soft clay layer in the range of 12 to 18 m
A-19
0
2
4
6
8
10
12
14
16
18
20
00 02 04 06 08 10 12D
ep
th (
m)
CPTu qt and u2 (kPa)
sand (SP)
Blodgett
Park Ridge
Deerfieldsoft clay
ClayLayerof Study
Northwestern UnivNational Test Site
Figure A12 Results of CPTu sounding at the NGES at Northwestern University Illinois
Post-processing of the CPTu data using Equations A7 A8 and A9 show good agreement in evaluating the
profile of effective yield stress in this clay layer compared with results from one-dimensional
consolidation tests made on undisturbed samples taken at the site
A-20
10
11
12
13
14
15
16
17
18
19
20
0 100 200 300 400 500 600 700 800 900 1000
De
pth
(m
)Effective Yield Stress sp (kPa)
Consols
033 qnet
053 Du2
060 qE
svo
Figure A13 Evaluation of effective yield stresses at NWU using consolidation tests and CPTu
A142 Application to soft Bay Mud California
Another example of a well-behaved fine-grained soil is the San Francisco Bay Mud which is a soft clay
deposit in northern California Results of a representative CPTu are shown in Figure A14 and indicate the
soil profile at a site in eastern San Francisco (Hunt et al 2002) For this site index values include 70 lt
wn lt 90 60 lt LL lt 80 and 35 lt PI lt 45 Post-processing the CPTu data using Equations A7 A8 and
A9 show good agreement with each other as well as with the results of consolidation tests as seen in
A-21
Figure A15 The consolidation tests were performed using a CRS device as reported by Pestana et al
(2002)
San Francisco Bay Mud (Pestana et al 2002)
0
5
10
15
20
25
30
35
40
0 1000 2000 3000
De
pth
(m
ete
rs)
Cone Resistance qt (kPa)
0
5
10
15
20
25
30
35
40
0 10 20 30 40 50 60
Sleeve Friction fs (kPa)
0
5
10
15
20
25
30
35
40
0 500 1000 1500
Porewater u2 (kPa)
Miscellaneous Fill
Young Bay Mud
Young Bay Mud
Clayey Sand
Miscellaneous Fill
Young Bay Mud
Young Bay Mud
Figure A14 Representative CPTu in San Francisco Bay Mud (data from Hunt et al 2002)
A-22
San Francisco Bay Mud (Pestana et al 2002)
0
5
10
15
20
25
30
0 100 200 300 400 500 600
De
pth
(m
ete
rs)
Preconsolidation Stress sp (kPa)
svo
033 qnet
053 Du2
060 qE
CRS
svo
Figure A15 Evaluation of effective yield stresses in Bay Mud from consolidation tests and CPTu
A15 CPTu Screening of Soft Organic Clays
For soft organic clays the Equations A7 A8 and A8 will not apply thus can serve as a means to
screen such geomaterials from the soil profile Two examples of CPTu soundings in soft organic
clays are presented to illustrate the approach Additional case records and documentation of
CPTu data in organic clays are given in Agaiby (2018)
A151 Soft organic alluvial soils in Washington DC
Results of in-situ tests including CPTu soundings in soft organic clayey silts along the Potomac River at the
Anacostia Naval Air Station are presented by Mayne (1987) Here the soft soils are categorized as organic
clayey silts (OH) per the Unified Soils Classification Systems (USCS) with mean values (and plus and minus
A-23
one standard deviations) of wn = 68 plusmn 16 LL = 83 plusmn 25 and PI = 37 plusmn 17 A representative
piezocone sounding is depicted in Figure A16 For the post-processing of CPTu data the use of the Q-F
and Q-Bq graphs do not find any points within the Zone 2 category for organic soils When the data are
evaluated to determine the profile of effective yield stress the Equations A7 A8 and A9 do not agree as
shown by Figure A17
0
5
10
15
20
25
0 2000 4000 6000
Dep
th (
met
ers)
Cone Resistance qt (kPa)
0
5
10
15
20
25
0 20 40 60 80 100
Sleeve Friction fs (kPa)
0
5
10
15
20
25
0 200 400 600
Pore Pressure u2 (kPa)
Figure A16 Representative CPTu in soft organic clay along the Potomac River DC
A-24
0
5
10
15
20
25
0 3 6 9D
epth
(m
eter
s)Soil Behaviour Type Zone
Q and Fchart
0
5
10
15
20
25
0 100 200 300 400 500 600
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
033qnet
053 Du2
06 qE
Oed
Figure A17 Post-processing of CPTu data from Anacostia-Bolling (a) SBTn zones using CPT material
index Ic (b) CPTu screening equations for yield stress
Note that the hierarchy of the results from Equations A7 A8 and A9 show that the CPTu
evaluations for preconsolidation stress in soft organic clays indicates
[A10] σp = 053 Δu2 lt 033 qnet lt 060 qE
A152 Soft organic peaty clay in Saint Paul MN
Results of CPTu tests were collected during a training exercise underneath I-35E just northeast
of Saint Paul MN in 2007 All three MnDOT CPT rigs available at the time were used to obtain
in-situ test data at the site The site is well known as having soft organic soils and samples were
obtained from three borings made at the site (Boring ID T523 T542 and T518) Boring logs
A-25
indicate the presence of highly organic silt loams with marl and spongy organic clayey silts
From 12 samples the natural water contents ranged from 30 to 191 with a mean of 112 plusmn
58
A representative piezocone sounding from the site (MnDOT CPTu ID F22Y0703C) is used for
this illustration and is presented as Figure A18 Post-processing these data using the SBTn
algorithms and CPT material index show that the soils classify primarily as Zone 3 (clays) and
Zone 4 (silty mix) thus do not identify the geomaterials properly as Zone 2 (organic soils)
0
5
10
15
0 500 1000 1500 2000
Dep
th (
met
ers)
Cone Tip Resistance
qt (kPa)
0
5
10
15
0 10 20 30 40 50 60
Sleeve Friction
fs (kPa)
0
5
10
15
-100 0 100 200 300 400
Porewater Pressure
u2 (kPa)
Figure A18 MnDOT CPTu sounding at the I-35E test site near St Paul Minnesota
A-26
1
10
100
1000
01 1 10
Norm
aliz
ed
Con
e T
ip R
esis
tan
ce
Q
Normalized Friction Fr = 100 fst(qt - svo) ()
9-Zone Soil Behavioral Type Chart
St PaulGravelly Sands(zone 7)
Sands
(zone 6)
Sandy Mixtures(zone 5)
Silt Mix(zone 4)
Clays
(zone 3)
Organic Soils(zone 2)
Sensitive Clays
and Silts(zone 1)
Very stiff OC clayto silt (zone 9)
Very stiffOC sandto clayey
sand(zone 8)
1
10
100
1000
-06 -04 -02 00 02 04 06 08 10 12 14
No
rmal
ized
Co
ne
Tip
Res
ista
nce
Q
Normalized Porewater Bq = Du2(qt-svo)
Clays(zone 3)
Sensitive(zone 1)Organic
(zone 2)
Gravelly Sands(zone 7)
Sands(zone 6)
Figure A19 Post-processing St Paul sounding using the 9-zone SBTn classification system
Using the recommended approach Figure A20 shows the CPT-evaluated yield stress profiles from
equations A7 A8 and A9 Here again the three estimates of σp do not agree but show the same
hierarchy as detailed in Equation A10
A-27
0
5
10
15
20
0 50 100 150 200 250 300 350 400
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
06 qeff svo
033qnet Consols
054 Du2
Figure A20 CPTu screening of soft organic soils at St Paul test site
A16 CPTu Evaluations of Yield Stress in Soft Organic Soils
Once the presence of soft organic clays and silts is identified using the aforementioned CPT
screening algorithms the evaluation of effective yield stress is conducted using the procedure
outlined in Chapter 2 of this MnDOT manual For soft organic soils an exponent of m = 090 is
recommended (Mayne et al 2009) Therefore the preconsolidation stress is obtained from
(units of kPa)
prime = A11 120590119901 033 ∙ (119902119899119890119905)090
The algorithm can be expressed in dimensionless form by
prime = A12 120590119901 033 ∙ (119902119905 minus 120590119907119900)119898prime ∙ (120590119886119905119898100)1minus119898prime
A-28
where σatm = reference stress (= 100 kPa = 147 psi) and m = 090 for soft organic silts and clays
The approach is applied to the two former example case studies in Figure A21 (Washington DC)
and Figure A22 (St Paul MN) respectively showing good agreement with benchmark results
taken from laboratory consolidation tests on undisturbed samples of these sites in both cases
0
5
10
15
20
25
0 100 200 300 400 500 600
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
033qnet^090
Oed
Figure A21 Profiles of yield stress from consolidation tests and CPTu method for organic clayey silts at
Anacostia-Bolling site in Washington DC
A-29
0
5
10
15
20
0 20 40 60 80 100 120 140 160
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
CPT
Consols
120590119901prime = 033 119902119899119890119905
090 in kPa
Figure A22 Profiles of yield stress from consolidation tests and CPTu method for organic clayey silts at
MnDOT test site near Saint Paul Minnesota
A17 Soil Unit Weight
As shown by Figure A23 total soil unit weight (t) can be estimated from the CPT sleeve friction resistance
(Mayne 2014) The equation can be expressed by
120574119905 = 120574119908 ∙ [122 + 015 ∙ 119897119899 (100 lowast 119891119904120590119886119905119898 + 001)] A13
where w = unit weight of water and satm = atmospheric pressure
A-30
Figure A23 Evaluation of soil unit weight from CPT sleeve friction
A18 Effective Stress Friction Angle
Sands
The effective stress friction angle () is one of the most important soil properties as it governs the strength
of geomaterials as well as affects soil-pile interface and pile side friction While an effective cohesion
intercept (c) can also be considered this is usually reserved for cemented or bonded geomaterials or
unsaturated soils and may lose its magnitude with time ageing or with prolonged environmental
changes
For clean quartz to silica sands where porewater pressures are essentially hydrostatic (Bq = 0) the
following expression has been calibrated with triaxial compression test results from undisturbed sand
samples and normalized cone resistances as presented in Figure A24 (Mayne 2007 2014)
A-31
120601prime(119889119890119892) = 176deg + 110deg 119897119900119892 (1199021199051) A14
Figure A24 Evaluation of effective stress friction angle in quartz-silica sands from CPT
Note Relationship applies to drained soil behavior when Ic lt 26 andor Bq lt 01
where qt1 is an earlier form of stress normalized cone tip resistance for sands given by
(119902119905 120590119886119905119898) 1199021199051 = A151 05 prime (120590119907119900120590119886119905119898)
A-32
In terms of units of bars this is expressed simply as (in units of bars)
119902119905 1199021199051 = 05 A152 prime ) (120590119907119900
Recently Robertson amp Cabal (2015) recommended the use of the modified form of normalized cone
resistance (Qtn) in Equation A10
120601prime (119889119890119892) = 176deg + 110deg 119897119900119892 (119876119905119899) A16
In sands qnet asymp qt since the overburden term is small relative to the cone tip resistance Figure A25 shows
that the two normalizations give very comparable results
Figure A25 Comparable values from qt1 and Qtn normalization schemes for CPTs in sands
Clays and Silts
A-33
In the case of soft to firm intact clays and silty clays the effective friction angles are determined from the
normalized cone resistance and porewater pressure parameters (Senneset et al 1989 Mayne 2016) as
shown in Figure A26 The exact solution when the angle of plastification = 0 is given as
qBQ
)tan1(tan61
1)tanexp()245(tan 2
A17
Approximate NTH Solution for from CPTu
qBQ
)tan1(tan61
1)tanexp()245(tan2
Approx asymp 295deg∙Bq0121∙[0256 + 0336∙Bq + log Q ]
1
10
100
20 25 30 35 40 45
Co
ne
Res
ista
nce
Nu
mb
er Q
(degrees)
Bq
010204060810
c = 0 = 0deg
Theory (dots)
Approximation (lines)
Theory
Figure A26 Evaluation of effective stress friction angle in clays and silts from CPTu results
Note Relationship applies to undrained soil behavior when Ic ge 26 andor Bq ge 01
which can be approximately inverted into the form (Mayne 2007)
A-34
0121 120601prime = 295deg ∙ 119861119902 ∙ [ 0256 + 0336 ∙ 119861119902 + log(119876)] A18
This algorithm is specifically applicable for the following ranges of porewater pressure parameter (01 le
Bq lt 1) and effective stress friction angles (20deg le lt 45deg)
A19 Stress History
The stress history can be characterized by an apparent yield stress or preconsolidation stress (sp) as well
as by its normalized and dimensionless form YSR = spsvo = yield stress ratio The YSR is in effect the
same as overconsolidation ratio (OCR) however now generalized to accommodate mechanisms of
preconsolidation that occur beyond just erosion glaciation and removal of overburden stresses but also
due to ageing desiccation repeated cycles of wetting-drying bonding repeated freeze-thaw cycles
groundwater changes and other factors
A-35
A generalized approach for evaluating the yield stress or preconsolidation in soils using net cone
resistance has been formulated as presented in Figure A27 (Mayne 2015) Yield stress in soils from CPT
10
100
1000
10000
10 100 1000 10000 100000
Yie
ld S
tre
ss
s
y
(kP
a)
Net Cone Resistance qt - svo (kPa)
Blessington
General Trend
sy = 033(qt-svo)m
Fissured Clays m = 11
Intact clays m = 10 Sensitive clays m = 09
Silt Mixtures m = 085Silty Sands m = 080
Clean Sands m = 072
m = 11 10 09
085
080
072
Units kPa
Fissured Clays
Figure A27 Evaluation of yield stress or preconsolidation stress in soils from CPT
The algorithm can be expressed in dimensionless form by
1minus119898prime prime 120590119886119905119898 120590119901 = 033 ∙ (119902119905 minus 120590119907119900)119898prime
( ) A19 100
where m = exponent depends on soil type m = 1 (intact clays) 085 (silts) 080 (sandy silts to silty sands)
and 072 (sands) For fissured clays the exponent m may be 11 or higher depending upon the age of the
A-36
formation and degree of jointing and discontinuities Note that fissured clays can be identified when Ic lt
26 and Bq lt 01
The exponent m has also been calibrated with CPT material index as presented in Figure A28
An algorithm to express this relationship for non-fissured soils is given by
25)652(1
2801
cIm
A20
This allows the post-processing of CPT to automatically choose the appropriate exponent m for
a layer by layer analysis
Figure A28 Yield stress exponent m in terms of CPT material index Ic
A-37
A110 Lateral Stress Coefficient
The horizontal geostatic state of stress is represented by the lateral stress coefficient K0 = shosvo
commonly referred to as the at-rest condition The magnitude of K0 for soils that have been loaded and
unloaded can be approximately estimated from
119870119900 = (1 minus 119904119894119899120601prime) ∙ 119874119862119877119904119894119899120601prime A21
Data compiled from in-situ K0 measurements using self-boring pressuremeter tests (SBPMT) total stress
cells (TSC) or push-in spade cells andor laboratory methods (instrumented consolidometers triaxials
andor suction measurements) on a variety of clays silts and sands have been compiled and reported by
Ku amp Mayne (2015) as shown in Figure A29 verifying Equation A15 as a means for evaluating K0 in soils
Figure A29 Relationship between lateral stress coefficient K0 and YSR or OCR in soils (Ku amp Mayne
2015)
A-38
A111 Undrained Shear Strength
The loading of soils can occur under conditions of being fully drained (Du = 0) partially drained or full
undrained (DVV0) where Du = excess porewater pressures (above hydrostatic) and DVV0 = volumetric
strain Further details on specific stress paths are best explained in terms of critical state soil mechanics
or CSSM (Mayne et al 2009 Holtz Kovacs amp Sheahan 2011) The prevailing drainage conditions depend
upon the rate of loading and permeability characteristics of the soil Normally in sands that are pervious
and exhibit high permeability a drained response occurs Exception to this may occur in loose sands
during fast earthquake loading resulting in soil liquefaction In clays that exhibit low permeability a fast
rate of loading will result in undrained loading at constant volume This in fact is a temporary and
transient condition often termed short term loading In the long term eventually porewater pressures
will dissipate (albeit slowly) and a drained response will prevail thus termed long term loading
The overall soil strength is controlled by the effective stress strength envelope Most commonly this is
represented by a simple linear relationship termed the Mohr-Coulomb criterion where the maximum
shear stress (max called the shear strength) is given as
= 119888prime + 120590prime ∙ 119905119886119899 120601prime A22 120591119898119886119909
where c = effective cohesion intercept s = effective normal stress and = effective stress friction angle
As a starting point values of c = 0 and = 30deg can be adopted for all soil types at least until the specific
soil type geologic formation and results from high-quality laboratory or field test data are available
Peak Undrained Shear Strength from Stress History
Using simplified critical state soil mechanics the peak undrained shear strength (max = su) can be
evaluated from the effective stress strength envelope (c = 0 effective friction angle ) and stress history
(ie OCR) in the form
prime DSS 119904119906 = (119904119894119899120601prime2) ∙ (119874119862119877)Λ ∙ 120590119907119900 A23
A-39
which corresponds to a simple shear mode Figure A30 presents the aforementioned relationship
together with data from 17 different clays tested in simple shear
Figure A30 Normalized undrained shear strength from simple shear tests on various clays
versus YSR or OCR
For the triaxial compression (TC) mode the equation would give slightly higher strengths that are
calculated from
prime TC 119904119906 = (1198721198882) ∙ (1198741198621198772)Λ ∙ 120590119907119900 A24
where Mc = (6 middot sin)(3-sin)
Peak Undrained Shear Strength from CPTu
A more direct approach to assessing undrained shear strength is via bearing capacity theory
whereby
A-40
A25 kt
votu
N
qs
s
where Nkt = bearing factor that depends on mode of shearing sensitivity of the clay degree of
overconsolidation and other factors For soft offshore clays the back figured Nkt from data collected at
14 well-documented sites (Low et al 2010) determined mean values based on mode of shearing Nkt =
119 (triaxial compression) Nkt = 136 (simple shear) and Nkt = 133 (vane)
A recent study of 51 clays that were tested by both field piezocone and laboratory CAUC triaxial tests
showed that essentially Nkt = 12 for intact clays and clayey silts of low to medium sensitivity (Mayne
Peuchen amp Baltoukas 2015) Figure A31 shows the slightly different trends for offshore versus onshore
clays For sensitive clays a lower value of Nkt = 10 would be appropriate and for fissured clays a higher
value of Nkt would be in the range of 20 to 30
A-41
A-42
Figure A31 Database from 51 clays with CPT qnet versus lab measured triaxial shear strength
(Mayne Peuchen amp Baltoukas 2015)
For soft to firm clays an alternate means to evaluate undrained shear strength is via the excess porewater
pressures ( u = u2 - u0) from the following
u
uN
us
D
D 2
A26
where N u = porewater bearing factor also dependent on the aforementioned factors The study by
Low et al (2010) found N u = 59 (triaxial compression) N u = 69 (simple shear) and N u = 71 (vane
shear)
A third alternate is found from the effective cone resistance (qE = qt - u2) whereby
kE
tu
N
uqs 2
A27
where NkE is a bearing term for this approach Mayne et al (2015) indicated a mean value of NkE = 8 for
soft-firm intact clays
Remolded Undrained Shear Strength from CPT
The remolded undrained shear strength (sur) is obtained from either field vane tests lab mini-vane shear
tests or lab fall cone devices This affords the evaluation of the clay sensitivity (St) which is defined as the
ratio of peak to remolded strengths at a given water content
119904119906(119901119890119886119896) 119878119905 = frasl A28 119904119906119903
For the CPT the sleeve friction has been noted to give values that are comparable to the
remolded undrained shear strength (Powell amp Lunne 2015) Therefore
asymp 119891119904 A29 119904119906119903
A112 Ground Stiffness and Soil Moduli
The stiffness of the ground can be represented by several geoparameters including the compressibility
indices (Cr Cc Cs) spring constants (kv) and moduli Regarding the latter there are several moduli that
are used in geotechnical engineering including shear modulus (G and Gu) Youngs modulus (E and Eu)
constrained modulus (D) bulk modulus (K) resilient modulus (MR) and subgrade reaction modulus (ks)
A-43
Soil Modulus
The definition of modulus is taken as E = DsD with Figure A32 showing a typical deviator stress (q = s1
- s3) versus axial strain (a) curve Theoretical interrelationships between the elastic moduli G E D and
K have a dependence on the Poissons ratio v The value of Poissons ratio can be taken as vu = 05 for
undrained loading (ie constant volume) while for drained loading which is accompanied by volumetric
strains a value of v = 02 may be used If we adopt the reference modulus as E then the
interrelationships with the other elastic moduli are given by
a Reference Stiffness 119864prime = drained Youngs modulus A30
119864prime
b Shear Modulus 119866prime = A31 [2(1+120584prime)]
119864prime (1minus120584prime) c Constrained Modulus 119863prime = A32
[(1+120584prime)(1minus2120584prime)]
119864prime
d Bulk Modulus 119870prime = A33 [3middot(1minus2120584prime)]
A-44
Figure A32 Representative stress-strain-strength curve for soils in triaxial compression
For the extreme case when = 0 in fact D = E When = 02 the two moduli are only 10
different D = 11middotE thus the constrained modulus and drained Youngs modulus are often
considered somewhat interchangeable
The resilient modulus (MR) applies to pavement analysis and design most commonly measured by cyclic
triaxial testing under repeated load applications In fact MR is a special case of Youngs modulus that
relates to the small strain stiffness measured in the nondestructive range but has developed permanent
plastic strains after many cycles of loading (Brown 1996 Dehler amp Labuz 2007)
The subgrade reaction modulus is actually a combined soil-structural parameter as its value
depends on the ground stiffness and the size of the loaded element The subgrade modulus is
defined as ks = q where q = applied stress and = measured deflection In terms of elasticity
solutions the deflection of a flexible circle of diameter d is given by = qmiddotdmiddot(1-2)E Therefore
119864prime
119896119904 = A34 [119889middot(1minus1205842)]
which has units of kNm3 or pcf
Table A2 summarizes several selected studies towards the approximate evaluation of these moduli
obtained directly from measured CPT data Note however that the results from in-situ penetrometer
tests generally represent a peak strength as indicated in Figure A32 Thus the measured cone tip
resistance (qt) reflects the top of the stress-strain curve either the undrained strength in clays (su) or the
effective friction angle () of sands or even a strength intermediate between these two values Thus
A-45
Source Soils Studied Reference Modulus
Findings
Kulhawy amp Mayne 1990
8 stiff OC clays Lab constrained modulus
D asymp 825 middot (qt - svo)
Mohammed et al (2000)
Resilient modulus
Abu-Farsakh 2004
7 Louisiana sites
Lab constrained modulus
D asymp 358 middot (qt - svo)
Mayne (2007b)
All soil types Constrained moduli from lab consolidation
First-order estimate D asymp middot(qt - svo)
Robertson (2009)
All soil types Constrained modulus from consolidation
D = m middot (qt - svo)
when Ic gt 22
1 m = Qtn when Qtn le 14
2 m = 14 when Qtn gt 14
when Ic lt 22 then
= 003 middot 10055 Ic + 168 m
Liu et al (2016)
16 clay sites in China
Resilient modulus MR = (146qt 053 + 1355 fs
14 + 236)244
(all in MPa)
Casey et al (2016)
8 clays Triaxial tests for undrained Youngs modulus
Eu(svc)07 = 316 - 23 middot LL
where Eu and svc (MPa)
and LL = liquid limit ()
qt is probably not the best means to obtain the slope of the stress-strain curve Instead the SCPTu that
provides the initial tangent shear modulus (Gmax) from the shear wave velocity (Vs) is a better choice
Table A2 Selected Modulus Relationships with Cone Penetration Test Measurements
A-46
Nonlinear Modulus
The small-strain shear modulus (Gmax or G0) represents the initial stiffness of all soils and rocks It is the
beginning of all stress-strain-strength curves for geomaterials and obtained from elasticity theory
119866119898119886119909 = 1198660 = 120588119905 middot 1198811199042 A35
where Vs = shear wave velocity t = tga = total soil mass density t = total soil unit weight and ga =
gravitational acceleration constant (98 ms2)
From a general viewpoint on stiffness the shear modulus G of soil is defined as the slope of shear stress
versus shear strain G = DDs for a tangent definition and by G = s as a secant definition The shear
modulus is related to its associated Youngs modulus E = 2G(1+v) Both moduli are in fact highly
nonlinear ranging from a maximum value at the small-strain stiffness (Gmax = G0 = t middotVs2 where t =
total soil mass density and Vs = shear wave velocity) to intermediate G values at medium strains (asymp 1)
to low values at peak strength As such a variety of algorithms and formulae have been developed to
represent either a partial range or the full stress-strain-strength behavior of soils over a range of
interests (eg Mayne 2005) In these formulations a variable number of input parameters may be
required in order to produce a stress-strain-strength curve
A modified hyperbolic form suggested by Fahey amp Carter (1993) has favorable attributes in that the
modulus reduction factor can be established with only a single variable This allows the initial stiffness
to the be small-strain stiffness (G0) that is reduced in terms of level of mobilized strength eg max =
qqmax = 1FS which is simply the reciprocal of the calculated factor of safety (FS) The magnitude of
secant shear modulus (G) corresponding to the particular level of loading is given by
119866 = 119872119877119865 middot 1198660 = (1198661198660) middot 1198660 A36
A-47
where the MRF = modulus reduction factor determined as
1 minus (120591 )119892 119872119877119865 = (1198661198660) = frasl A37 120591119898119886119909
and g = empirical exponent term Figure A33 shows the relationships for normalized shear stress ()
versus shear strain (s) and corresponding normalized secant shear modulus (G = s) with shear strain
over a range of exponent values 01 le g le 10 For a discussion of tangent G please see Fahey amp Carter
(1993)
Using laboratory data from resonant column-torsional shear tests and triaxial specimens with local
strain measurements for Gmax reference values modulus reduction curves for a selection of sands and
clays are presented in Figure A34 The data indicate that the exponent g falls within the range 02 lt g lt
05 for many soils tested drained and undrained A mean value of g = 03 is recommended for
preliminary site investigations and designs until additional information can be obtained
The value of max is the shear strength of the soil generally taken as either (1) drained (c = 0) max =
svo middot tan or (2) undrained where max = su = undrained shear strength as discussed earlier Thus each
shear stress () can be associated with its shear modulus (G) and the relevant shear strain is found from
s = G This allows for generation of nonlinear stress-strain-strength curves at all depths from SCPTu
data in clays sands and mixed soil types
A-48
Modified Hyperbola (Fahey amp Carter 1993 Canadian Geotech J) Coefficient Term f = 1
``
0
01
02
03
04
05
06
07
08
09
1
000001 00001 0001 001 01 1
No
rmal
ized
GG
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
0
01
02
03
04
05
06
07
08
09
1
0 01 02 03 04 05 06 07 08 09 1
No
rmali
zed
GG
ma
x
Mobilized Shear Stress max
g =1
07
05
03
02
01
0
01
02
03
04
05
06
07
08
09
1
00001 0001 001 01 1
No
rmali
zed
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
0
01
02
03
04
05
06
07
08
09
1
0 1 2 3 4 5 6 7 8 9 10
No
rmali
zed
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
Reduction Factor
GGmax = 1- (max)g
Figure A33 Normalized modulus (GGmax) and normalized shear stress (max) versus shear strain (s)
for modified hyperbolic relationship of Fahey amp Carter (1993)
A-49
0
01
02
03
04
05
06
07
08
09
1
0 01 02 03 04 05 06 07 08 09 1
Mo
du
lus
Re
du
cti
on
G
Gm
ax
or
EE
ma
x
Mobilized Strength max or qqmax
NC SLB Sand
OC SLB Sand
Hamaoka Sand (e=0625)
Hamaoka Sand
Toyoura Sand e = 067
Toyoura Sand e = 083
Ham River Sand
Ticino Sand
Kentucky Clayey Sand
Kaolin
Fujinomori Clay
Pietrafitta Clay
London Clay
Vallericca Clay
Pisa Clay
Pisa Clay
Pisa Clay
Pisa Clay
Kiyohoro Silty Clay
Thanet Clay
Exp g = 02
Exp g = 03
Exp g = 04
Exp g = 05
MRF = 1 - (qqmax)g
Figure A34 Modulus Reduction Factors (MRF) with mobilized strength of clays and sands
A113 Coefficient of Consolidation
The rate at which foundation and embankment settlements occur as well as the dissipation of excess
porewater pressures is controlled by the coefficient of consolidation (cv) The magnitude of cv is also
required in the design of vertical wick drains that can be installed in soft ground to expedite the time for
consolidation Using the results of CPTu dissipation tests which measure the rate at which the u2 readings
vary with time the in-situ profile of cv can be evaluated Often the piezocone-interpreted values of cv
are validated by comparison with results from laboratory one-dimensional consolidation tests (eg Abu-
Farsakh MY 2004) Alternatively a better method is to cross-check the values with the measured full-
scale performance of instrumented embankments that are constructed over the ground and the recorded
time-rate of consolidation and settlements can provide the best cv for that site (Abu-Farsakh MY et al
2011)
A-50
Monotonic Dissipation Tests
An illustrative dissipation curve from piezocone testing at IDT State Highway 95 at an embankment and
bridge crossing is presented in Figure A35 After the CPTu sounding was halted at a depth of 512 feet
the recorded decay of porewater pressures was observed to be monotonic with time until the dissipations
were ended at 1000 seconds During that time the u2 readings decreased from 115 psi to 47 psi At this
location the depth to the groundwater table is about 16 feet thus the calculated equilibrium water
pressure is 15 psi Commonly a characteristic time for piezo-dissipations is taken at 50 degree of
consolidation although other degrees may be adopted By evaluating the value of u2 at 50 as shown in
Figure A35 the characteristic t50 = 404 s can be obtained
0
20
40
60
80
100
120
1 10 100 1000 10000
Pore
wat
er P
ress
ure
u2
(psi
)
Time (s)
CPTU-02E-1 z = 1560 m
uo (hydrostatic)
u2 (initial) = 115 psi
u2 (final projected) = u0 = 15 psi
u2 at 50 = frac12(115+15) = 65 psi
Idaho DOT State Route 95Dissipation at 512 feet
t50 = 404 seconds
Time to reach50 consolidation
A-51
Figure A35 Piezo-dissipation at Sandpoint Bridge Crossing Idaho for depth z = 512 feet illustrating the
evaluation of characteristic time t50
It is also common to plot normalized excess porewater pressures relative to the measured initial Du2
that is obtained during penetration at the constant rate of 20 mms ie DuDui versus time Figure A36
shows a set of piezo-dissipation records for a bridge and embankment site in southern Louisiana
reported by the LTRC While the porewater pressure axis is arithmetic the time scale is often plotted on
either logarithmic or square root scales In either case the t50 is then found from the measured
dissipation curve when DuDui = 050
A-52
Figure A36 Set of monotonic piezo-dissipations taken at various depths for Courtableau Bridge
Site on Louisana State Highway 103 (Abu-Farsakh et al 2011)
Dilatory Dissipation Curves
In a number of situations a monotonic type dissipation does not occur on the onset but instead a dilatory
type curve is observed Here after stopping the push of the penetrometer the recorded porewater
pressures initially begin to rise and eventually reach a peak value then afterwards follow a phase where
the Du values decrease to hydrostatic (Figure A37) A full solution is available for both cases (Burns amp
Mayne 2002) Dilatory responses are most often associated with overconsolidated clays and silts
although occasionally are observed in fine-grained soils with low OCRs
The selection of the characteristic t50 value may found by using a square root of time plot to find the initial
u2 value by projection of the post-peak dissipatory data back to the ordinate axis as shown by the example
in Figure A38 In this example the initial u2 = 480 kPa and then the same procedures in Figure A35may
apply
A-53
Figure A37 Monotonic and dilatory porewater dissipation responses in soils
(after Burns amp Mayne 1998)
Figure A38 Method for obtaining t50 from dilatory dissipation curves
(Schneider amp Hotstream 2010)
A-54
Interpretation of Dissipation Test Data
There are a number of different solutions available for the interpretation of the coefficient of
consolidation (cv) from the piezocone dissipation curves For monotonic type response the strain path
method (SPM) developed at Oxford University (Teh amp Houlsby 1991) is well-recognized while the Georgia
Tech solution by Burns amp Mayne (2002) is based on spherical cavity expansion and critical state soil
mechanics (SCE-CSSM) and handles both monotonic and dilatory curves For both solutions a simplified
procedure can be recommended that relies on the aforementioned t50 value obtained from the measured
field data as given in Table A3 The units for cv are in cm2s as shown or equivalent
Table A3 Recommended procedures for calculating cv from t50 obtained in dissipation tests
Method Equation for coefficient of
consolidation cv
RemarksNotes Eqn
No
SPM
(Teh amp
Houlsby
1991) 50
2)(2450
t
Iac
Rc
v
t50 = measured time to reach 50
dissipation
IR = Gsu = undrained rigidity index
G = shear modulus
su = undrained shear strength
ac = penetrometer radius (ac = 178
cm for 10-cm2 cone ac = 220 for a
15-cm2 size)
A32
SCE-CSSM
(Burns amp
Mayne 2002) 50
7502 )()(0300
t
Iac Rc
v
A33
A-55
Evaluating rigidity index
Both the SPM and SCE-CSSM solutions require an estimate of the in-situ rigidity index (IR = Gsu) of the
soil If the results of SCPTU are available Krage et al (2014) have derived an expression for IR that
depends upon the small-strain shear modulus net cone tip resistance and effective overburden stress
250750
max50
8111
vonet
Rq
GI
s A38
where consistent units are input for Gmax qnet and svo The value of Gmax is obtained from B29 using the
measured shear wave velocity (Vs) and unit weight evaluated from B7
If the shear wave velocity is not measured it can be estimated from the CPT data A number of
methods have been reviewed for CALTRANS by Wair et al (2012) including one that is applicable
to sands silts and clays of low sensitivity that are inorganic and uncemented
119881 (119898119904) = [101 middot 119897119900119892(119902119905) minus 114]167 middot (100 middot 119891119904119902119905)03 A39 119904
where qt is in units of kPa (Hegazy amp Mayne 1995) An alternative relationship is given by Robertson amp
Cabal (2015)
A114 Hydraulic Conductivity
The hydraulic conductivity is a parameter that expresses the flow characteristics of the soil In
geotechnics it is also called the coefficient of permeability (k) and has units of cms or feetday
Through consolidation theory the hydraulic conductivity relates directly to the coefficient of
consolidation
A-56
A40 D
ck wv
where w = unit weight of water and D = constrained modulus
Therefore one approach to evaluating k can be from the site-specific cv obtained from the
dissipation tests using an estimate of D from one of the relationships given in Table A2
An approach developed for soft normally-consolidated soils that uses t50 directly to assess the magnitude
of k is presented in
Figure A39 (Parez and Fauriel 1988) The mean trendline through the wedges of soil type can be
approximated by
251
50 (sec)251
1)(
tscmkh
A41
A-57
Figure A39 Hydraulic conductivity versus dissipation time for 50 consolidation
(Parez amp Fauriel 1988)
A-58
APPENDIX B
List of Figures
Figure B1 Conventional methods vs direct CPT approach to shallow foundation response B-3
Figure B2 Direct bearing capacity relationship for embedded square and strip footings situated on clean
sands (after Schmertmann 1978) B-6
Figure B3 Direct CPT bearing factor Rk as function of footing width B and embedment depth from finite
element analyses (Tand et al 1995) B-6
Figure B4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio and sand
consistency (Eslaamizaad amp Robertson 1996) B-7
Figure B5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp Salgado (2005) in
terms of footing size relative density and base settlement B-8
Figure B6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and normalized
cone tip resistance (after Eslami amp Gholami 2005) B-8
Figure B7 Direct relationship between ultimate bearing stress in clays and measured cone tip resistance
(Schmertmann 1978) B-10
Figure B8 Direct relationship between ultimate foundation bearing stress in clays and net cone tip
resistance (after Tand et al 1986) B-11
Figure B9 Measured load-displacements for each of the five large footings at TAMU (Briaud amp Gibbens
1994) sand site B-12
Figure B10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU (Briaud amp
Gibbens 1994) footings B-13
Figure B11 Characteristic stress versus square root normalized displacements for TAMU footingsB-15
Figure B12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sandB-16
Figure B13 Summary of sand formation factors with corresponding CPT resistancesB-19
Figure B14 Normalized foundation stresses vs square root of normalized displacement for spread
footings on sand (modified after Mayne et al 2012) B-20
Figure B15 Definitions of capacity from three types of observed foundation load test responses
(modified after Kulhawy 2004) B-21
Figure B16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footingsB-22
B-1
Figure B17 Measured load vs displacement curves for 3 shallow foundations on clay at Baytown Texas
(Stuedlein amp Holtz 2010)B-25
Figure B18 Characteristic stress vs square root of normalized displacement for all three foundations at
Baytown site (Stuedlein amp Holtz 2010) B-26
Figure B19 Normalized foundation stress vs square root of normalized displacements B-27
Figure B20 Applied foundation stress vs CPT-calculated stress for 67 footingsB-28
Figure B21 Applied footing stress vs CPT calculated stress on logarithmic scales B-29
Figure B22 Summary graph and equations of direct CPT method for evaluating the vertical B-30
Figure B23 Foundation soil formation parameter hs versus CPT material index Ic B-32
Figure B24 Bearing capacity ratio qmaxqnet versus CPT material index IcB-33
Figure B25 Influence factor for rectangular foundations from elastic theory solution (Mayne amp
Dasenbrock 2017) B-34
List of Tables
Table B1 Direct CPT methods for bearing capacity of footings on clean sandsB-4
Table B2 Direct CPT methods for bearing capacity of footings on clays B-9
Table B3 Summary of large footings soil conditions and reference sources of database B-17
Table B4 Sources for shallow foundation performace B-34
B-2
B1 DIRECT CPT METHOD FOR SHALLOW FOUNDATIONS
B11 Introduction
The analysis of shallow foundations on soils is classically handled as a two-step and two-part set of
calculations involving (a) bearing capacity commonly reliant on limit plasticity solutions and (b)
settlement or more appropriately termed displacements that are assessed via elastic continuum theory
The utilization of cone penetration tests (CPT) provides the necessary geotechnical data for the site-
specific information on the subsurface conditions at the project site including the geostratigraphy and
evaluation of input parameters This is still a viable approach where the CPT readings are interpreted to
give the soil unit weight (γt) effective friction angle (ϕ) undrained shear strength (su) preconsolidation
stress (σp) and elastic moduli (D and E) for analysis
An alternative approach is the use of CPT results to provide direct assessments of bearing capacity
andor settlements A comparison of these two distinctly different and alternate paths is depicted in
Figure B1
Figure B1 Conventional methods versus direct CPT approach to shallow foundation response
A number of available direct CPT approaches for shallow footings are reviewed within this report
Moreover as the entire load-displacement-capacity response of shallow foundations to loading occurs
as a continuous nonlinear phenomenon a single direct CPT solution is presented for the behavior of
B-3
footings on sands (drained) and clays (undrained) The methods discussed herein are specifically to
address the general case of vertical loading of shallow foundations Additional more complex situations
that consider load eccentricity moments inclined forces sloping ground and other facets may be
handled using well-established procedures that are discussed elsewhere (eg Vesić 1975 Kulhawy et
al 1983)
This report focuses on foundations on granular soils andor soils exhibiting drained behavior since
Federal Highway Administration (FHWA) studies have concluded that less than 1 of shallow
foundations for highway bridges are placed on clay soils (Paikowsky et al 2010) One reason for this is
because soft-firm intact clays require a more complicated process that assesses both a short-term
analysis (undrained) as well as long-term analysis (drained) thus requiring undrained bearing capacity
undrained distortion displacements drained bearing capacity and drained primary consolidation
settlements
B12 Direct CPT Methods for Bearing Capacity of Footings on Sands
Classical bearing capacity solutions are based most often in limit plasticity theory albeit can be
formulated from limit equilibrium cavity expansion and numerical methods Because the limit
plasticity solutions are prevalent and adopt total stress analysis they are usually developed as either
fully drained (Δu = 0) or fully undrained (ΔVV = 0) where Δu equals excess porewater pressure and
ΔVV equals volumetric strain Paikowski et al (2010) provides an extensive review of various solutions
There are a number of direct CPT methods for the evaluation of the bearing capacity of footings and
shallow foundations Table B1 lists a variety of these direct CPT approaches for footings on sands In
the direct approach a one-step process is used to scale the measured cone penetrometer readings (ie
measured cone tip resistance (qc) total cone tip resistance (qt) sleeve friction (fs) andor measured
porewater pressure u2)) to obtain the ultimate bearing stress (qult) of the foundation
Table B1 Direct CPT methods for bearing capacity of footings on clean sands
Method Surface Footing Remarks and Embedment Notes
Meyerhof (1956) qult = qc (B12) middotcw qult = qc (B12)(1+DeB)cw
Note for silty sand reduce by 05
cw = 10 dry or moist sand cw = 05 submerged sand
Meyerhof (1974) qult = qc (B + D)40 with stresses in tsf
Presumably dry sands B = fdn width (feet) and D = depth (feet)
Schmertmann (1978)
NA (applies to embedded footings)
Square qult =055σatm(qcσatm)078
Strip qult =036σatm(qcσatm)078
See Figure A2
Embedment applies De gt 05(1+B) for Blt1m De gt 12 m for B gt1m
B-4
Canadian qult = Rk0middotqc Applied to FS = 3 Geotech Society where Rk0 = 03 where FS = factor of (CFEM 199 2) safety
Tand et al (1995) NA qult = Rkmiddotqc + σvo See Figure A3 where Rk = fctn(De B)
Frank and Magnan (1995)
qult = Rk0middotqc where Rk0 = fctn(sand
consistency) Loose Rk0 = 014
Medium Rk0 = 011 Dense Rk0 = 008
qult = Rk1middotqc + σvo where Rk1 = function (De B L amp
sand consistency) Factor Rk1 = Rk0[1+Rk206+04(BL) (DeB)]
and Rk2 = 035 (loose) 050 (medium) and 085 (dense)
Loose sand qc lt 5 MPa
Medium sand 8 MPa lt qc lt 15MPa
Dense sand qc gt 20 MPa
Eslaamizaad amp Robertson (1996)
qult = KΦmiddotqc See Figure A4
See surface footing equation KΦ = function (BDe shape and density)
Lee amp Salgado (2005)
qbL = βbcmiddotqc(AVG)
where qc averaged over distance B
Not addressed See Figure A5 for factor βbc = fctn(B DR
K0 and sB)
beneath the base
Eslami amp Gholami (2005
2006)
See embedded footing solution
qult = Rk1middotqc where Rk1 = function (ratio DeB
and normalized qcσvo) See Figure A6
Measured qc and qcσvo are geometric means over 2B deep
beneath footing
Robertson amp Cabal (2007)
qult = KΦmiddotqc with KΦ = 016
See surface footing equation
Briaud (2007) qult = KΦmiddotqc with KΦ = 023
Based on full-scale tests at Texas AampM
Mayne amp Illingworth
(2010) qult = 018 qc-Mean
Note qc -Mean is averaged CPT cone resistance over depth of
influence z = 15 B deep
Based on 30 footing load tests on 12
sands
Lehane (2013) qult = 016 qc-AVE Note qc-AVE is averaged CPT cone resistance within z (m) = [B (m)
]07
Based on 47 load tests
Table B1 Continued
Notes B = minimum footing width (or diameter) L = footing length De = embedment depth cw = water
table correction σvo = total overburden stress at bearing elevation and σvo = the effective vertical stress
B-5
Figure B2 Direct bearing capacity relationship for embedded square and strip footings
situated on clean sands (after Schmertmann 1978)
Figure B3 Direct CPT bearing factor Rk as function of footing width B and embedment depth
from finite element analyses (Tand et al 1995)
B-6
Figure B4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio
and sand consistency (Eslaamizaad amp Robertson 1996)
B-7
Figure B5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp
Salgado (2005) in terms of footing size relative density and base settlement
Figure B6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and
normalized cone tip resistance (after Eslami amp Gholami 2005)
B13 Direct CPT Methods for Bearing Capacity of Footings on Clays
For direct CPT evaluations of foundation bearing capacity in clays Table B2 summarizes some of the
well-documented methods that are available Generally these assume that the loading is fast enough
and the permeability of the clay is also sufficiently low such that an undrained (ie constant volume)
condition is maintained This is fine for short term loading of foundations particularly for soft to firm
intact clays However the undrained case is not permanent Given sufficient time excess porewater
pressures in the clay will eventually dissipate and hydrostatic conditions will return to equilibrium
During this dissipation phase primary consolidation will occur that results in drained settlements Also
a drained bearing capacity condition will prevail As the CPT is advanced at a constant rate of 20 mms
the recorded readings normally constitute undrained behavior from a direct measurement viewpoint
Thus direct CPT methods are normally focused at an undrained foundation response Nevertheless the
drained footing case can be addressed using CPT results via use of conventional analysis approach and
effective stress limit plasticity solutions that require piezocone penetrometers with porewater pressure
measurements
B-8
B
D)
L
B4060(3501320kc
Many of the earlier solutions were based on data from mechanical penetrometers andor older electric
cones that measure the uncorrected tip resistance (qc) It is now well-recognized that this measurement
must be updated to the corrected cone resistance (qt) because of porewater pressure effects that act on
the mechanics of the load cell and geometry of the particular penetrometer design The results are
particularly significant in clays silts and mixed soils that are soft to firm to stiff where excess porewater
pressures are recorded
Table B2 Direct CPT methods for bearing capacity of footings on clays
Method Footing Comments NotesRemarks
Meyerhof (1974) qult = αbcmiddotqc
where 025 le αbc le 050
Applicable to saturated insensitive clays under short-term loading
Cone tip resistance
measured by
mechanical CPT
Trofimenkov (1974)
Approximation
qult asymp (qc33)09
(Assuming FS = 3)
Strip footings on clays and sandy clays 06m le B le 15m 1m le zemb le 25m
Mechanical CPT
with qc and qult in
kgcm2
Schmertmann (1978)
qult = function (qc and foundation shape)
Relationship shown in Figure A7 for square and strip footings
Cone tip resistance
measured by
mechanical CPT
Tand et al (1986)
qult = Rkmiddot(qc - σvo) + σvo
Factor Rk obtained from Figure A8 accounts for surface and deep footings Separate Rk factor for intact and jointed clays
= (qc1middotqc2)05 qc
where qc1 is geometric mean from bearing elevation to 05B deeper and qc2 is geometric mean from 05B to 15B beneath foundation base
Mix of data from
mechanical CPT qc
and electric CPT qc
LCPC Method
(Frank amp Magnan 1995)
Footing a surface
qult = kc middotqc + σvo
where kc = 032
Embedment case
Note Methods above generally consider undrained bearing capacity
B-9
Figure B7 Direct relationship between ultimate bearing stress in clays and measured cone tip
resistance (Schmertmann 1978)
B-10
Figure B8 Direct relationship between ultimate foundation bearing stress in clays and net
cone tip resistance (after Tand et al 1986)
B14 Direct CPT Assessments of Settlement and Displacements of Shallow Foundations
Methods for evaluating the magnitude of foundation displacements (s) can be found using
elastic theory subgrade reaction methods spring models and numerical simulations The
classical approach to footing settlement calculations on sands is to utilize elastic theory
solutions (eg Poulos amp Davis 1974) which take on the form
119902∙119861∙119868∙(1minus1205842) 119904 = B1
119864119904
where q = applied footing stress and I = displacement influence factor from elasticity theory
The value of I depends upon foundation geometry footing rigidity embedment depth variation of soil
modulus with depth compressible layer thickness and other factors as discussed by Mayne amp Poulos
(1999) For instance for a flexible circular footing of diameter B resting on an infinitely deep soil
formation having a constant modulus Es with depth the influence factor I is equal to 1 for the center
point The results of in-situ field tests such as pressuremeter tests (PMT) standard penetration tests
(SPT) cone penetration tests (CPT) flat dilatometer tests (DMT) and other methods can be used to
ascertain the input value of soil modulus Es
Alternatively direct CPT methods have been investigated to provide a one-step assessment of
foundation displacements Selected methods for direct CPT evaluation of shallow foundation
settlements on granular soils is include DeBeers amp Martens (1957) Meyerhof (1965) DeBeer (1965)
Thomas (1968) Schmertmann (1970) Meyerhof (1974) Berardi amp Jamiolkowski amp Lancellotta (1991)
Robertson (1991) Lutenegger amp DeGroot (1995) Lehane amp Dougherty amp Schneider (2008) Gavin amp
Adekunte amp OrsquoKelly (2009) Mayne amp Illingworth (2010) and Uzielli amp Mayne (2012)
B141 Large Scale Footing Load Tests
The measured load-displacement response of shallow foundations is conducted in the field using either
stepped loading procedures or continuous rate of displacement methods Results from the well-
documented test program at the Texas AampM University (TAMU) site (Briaud amp Gibbens 1999 Briaud
2007) for five spread footings on sand are shown in Figure B9 Each footing clearly shows a nonlinear
B-11
behavior to loading over the range of testing This site is one of the national geotechnical
experimentation sites (NGES) funded by the National Science Foundation (NSF) FHWA and American
Society of Civil Engineering (ASCE)
Figure B9 Measured load-displacements for each of the five large footings at TAMU (Briaud
amp Gibbens 1994) sand site
B142 Generalized Direct CPT Method for Footing Response on Soils
The use of applied foundation stress versus normalized displacement curves (q vs sB) is generalized to
footings on sands silts intact clays and fissured clays A square root plotting of the normalized
displacements (sB) permits a single parameter characterization for each specific soil type that in turn
correlates with the net cone tip resistance The generalization is based on a statistical review of data
from large foundations (B gt 05m) involving 70 full-scale load tests with available CPT data For fine-
grained soils the data are from electronic piezocone tests where the proper correction from qc to qt has
been obtained to allow the best possible results The methodology permits an evaluation of the load-
displacement-capacity response of shallow square footings based on CPTs performed in the four types
B-12
of ground conditions For the TAMU footings the characteristic stress-normalized displacement
response is shown in Figure B10 where all five foundations can be represented by a single curve
Figure B10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU
(Briaud amp Gibbens 1994) footings
B143 Characteristic Stress-Displacement Curves
Fellenius amp Altaee (1994) recommended the concept of a characteristic stress (q) vs normalized
displacement (sB) curve for foundation response on a given soil formation later supported by Decourt
(1999) Briaud amp Gibbens (1999) Lutenegger amp Adams (2003) and Briaud (2007) In this approach the
measured load (Q) vs displacement from individual footings of various sizes that rest on the same soil
conditions collapse to a unified relationship given in Equation B2
119939119943 119956 119954119938119953119953119949119946119942119941 = 119938119943 (119913
) B2
B-13
where af and bf are empirical fitting coefficients (Decourt 1999 Uzielli amp Mayne 2011 2012)
In a review of measured load tests data from large spread footings situated on different sands it has
been suggested that Equation B2 can be reduced to a single parameter expression by use of square root
plotting where the exponent bf is equal to 05 (Mayne amp Illingworth 2010 Mayne et al 2012)
119956 119954119938119953119953119949119946119942119941 = 119955119956radic
119913 B3
where rs = fitted soil parameter from best fit line regression In the case of sands the coefficient rs
represents the supporting soil conditions including particle size relative density and fines content as
well as geologic origin aging and overconsolidation effects
The results from the five TAMU footings are presented in this format in Figure B11 that shows the
formation factor rs = 486 MPa for this sand deposit This single coefficient can be used to express the
nonlinear load versus settlement of any size footing on this sand site Moreover one criterion from
Eurocode for bearing capacity that is used by the European community identifies that stress (or load)
which corresponds to (sB) = 10 That is when the displacement equals ten percent of the footing
width then the capacity has been reached Therefore adopting this criterion for footings on sands
the ultimate bearing stress (qult) for footings on sand can be taken when (sB) equals 010 or when
square root of (sB) equals 0316 In that case this gives qult equal to 0316 middot rs For the TAMU site this
gives qult equal to 0316 times (486) which equals 153 MPa as illustrated by Figure B12
B-14
Figure B11 Characteristic stress versus square root normalized displacements for TAMU
footings
B-15
Figure B12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sand
B15 Footing Database
A database of spread footings and large plates was assembled which considers only full-size shallow
foundations (05 le B le 6 m) that rest on sands silts and clays Sites for the foundations were also subjected to CPT The inclusion solely of large footings are significant because results from small-scale
model footings exhibit scale effects In fact experimental centrifuge work and numerical simulation
studies have shown that bearing capacity factors are size-dependent especially for factor Nγ decreasing
with footing width B (Kimura et al 1985 Cerato amp Lutenegger 2007 Mase amp Hashiguchi 2009) A
direct approach based on full-size footings provides a more reliable means for foundation evaluation
The compiled database is given in Table B3 with a total of 70 large footings and plates These include a
listing of 34 foundations on 13 sands 11 footings on 4 silt deposits 13 footings or large plate load tests
on 6 intact clays and 12 foundations on 5 fissured clays Most of the footings were square or nearly
square (80) while the remaining were circular The largest foundation was a mat 5 m by 14 m in plan
loaded by stacking Kentledge concrete blocks to cause a bearing failure in the underlying soft clay For
sands the largest footing consisted of the 6-m square concrete pad In terms of equivalent square
footings mean footing sizes were 155 m (sands) 114 m (silts) and 126 m (clays)
B-16
Additional details on the individual load tests and site conditions are given elsewhere (Mayne 2009
Mayne amp Illingworth 2010 Uzielli amp Mayne 2011 2012 Mayne et al 2012 Mayne amp Woeller 2014)
Table B3 Summary of large footings soil conditions and reference sources of database
Sand Site Location Soil Conditions Footings Numbers
Shapes and Sizes
ReferencesSource
College Station
Texas Pleistocene sand 5 Square 1 15 25 33 m Briaud amp Gibbens (1999)
Kolbyttemon Sweden Glaciofluvial sand 4 Rect B = 06 12 17 24 m
Bergdahl et al (1985 1986)
Fittija Sweden Glaciofluvial sand 3 Rect B = 06 17 24 m Bergdahl et al (1984 1985)
Alvin West Texas Alluvial sand 2 Circular D = 235 m Tand et al (1994)
Alvin East Texas Alluvial sand 2 Circular D = 22 m Tand et al (1994)
Perth Australia Silceous dune sand 4 Square B = 05 and 10 m
Lehane (2008)
Grabo T1C Sweden Compacted sand fill
1 Square B = 046 m Long (1993)
Grabo T2C Sweden Compacted sand fill
1 Square B = 063 m Long (1993)
Grabo T3C Sweden Compacted sand fill
1 Square B = 080 m Long (1993)
Labenne France Dune sand 6 Square B = 07 and 10 m
Amar et al (1998)
Green Cove Florida Brown silty sand 1 Circular D = 182 m Anderson et al (2006)
Durbin South Africa
White fine sand 1 Square B = 609 m Kantley (1965)
Porto Portugal Residual clayey sands
1 Circular D = 12 m and 1 plate
Viana da Fonseca (2003)
Jossigny France Soft clayey silt 2 Square B = 1 m Amar et al (1998)
Tornhill Sweden Glacial Baltic till 3 Square B = 05 1 2 m Larsson (2001)
Vagverkel Sweden Stiff medium silt 3 Square B = 05 1 2 m Larsson (1997)
Vattahammar Sweden Brn-Gry layered silt 3 Square B = 05 1 2 m Larsson (1997)
B-17
Table B3 Continued
Clay Site Location Soil Conditions Footings Numbers Shapes and Sizes
ReferencesSource
Baytown Texas Fissured Beaumont 2 plates with d = 076 m Stuedlein amp Holtz (2008)
Belfast Ireland Soft clay silt ldquosleechrdquo
1 Square Pad B = 20 m Lehane (Geot Engr 2003)
Bothkennar Scotland Soft silty clay 2 Square Pads B = 22 24 m
Jardine et al (1995)
Bangkok Thailand Soft to stiff clay 4 Square 067 075 090 10 m
Brand et al (1972)
Haga Norway Stiff OC clay 2 Square Footings B = 10 m
Andersen amp Stenhammar (1982)
Rio Grande Brazil Sandy residual clay 3 Square 04 07 and 10 m
Consoli et al (1998)
Shellhaven England Soft estuarine clay Rectangular 5 m by 14 m Schnaid et al (1993)
Texas City A Texas Coast Fissured Beaumont 3 circular plates d = 058 m
Tand et al (1986)
Texas City B1 Texas Coast Fissured Beaumont 2 circular plates d = 058 m
Tand et al (1986)
Texas City B2 Texas Coast Fissured Beaumont 1 circular plate d = 058 m
Tand et al (1986)
Alvin Texas Texas Coast Fissured Beaumont 3 circular plates d = 058 m
Tand et al (1986)
For the series of footings on sands Figure B13 shows the summary of measure formation factors (rs) for
the 34 foundation load tests Also indicated on the graph are the corresponding mean qc values of the
sands averaged over 15middotB beneath the bearing elevations
Note also in sands that the qt and net resistance (qnet) are all very close to the measured qc since (a)
porewater pressure corrections for the a-net value are negligible in clean sands and (b) overburden
stresses are typically only 1 to 3 of the magnitude of qc This is especially true of shallow depths
Therefore for clean sands qnet is approximately qt which is approximately qc
B-18
Figure B13 Summary of sand formation factors with corresponding CPT resistances
As suggested by Briaud (2007) various in-situ measurements can be used to normalize the results from
cone penetration tests in this case Herein the qc from the CPT soundings in the sands are used to
normalize the footing stress axis as presented in Figure B14 It is evident that the results of the load-
displacement response of all 34 footings on 13 different sands can be captured by the characteristic
stress versus normalized displacement with the inclusion of the site-specific cone tip resistance In this
case the mean trend for all footings and sand sites indicates a simple expression shown in Equation B4
119954119943119952119952119957119946119951119944 119956 Clean quartz-silica sands = 120782 120787120790120787radic B4
119954119940 119913
B-19
Figure B14 Normalized foundation stresses vs square root of normalized displacement for
spread footings on sand (modified after Mayne et al 2012)
The results of measured force-displacement curves (Q vs s) from footing load tests are nonlinear and
thus the definition of capacity must be addressed Kulhawy (2004) identified three main curve types
that occur during load testing depicted in Figure B15 The most common response (Type A) shows no
clear peak and is observed in sands and silts as well as slow loading of insensitive clays and fissured
clays In this case ldquocapacityrdquo can be defined by the Eurocode corresponding to the load (or stress) when
sB=10 (Amar et al 1998) For foundations situated on many clays a plateau capacity is reached as
depicted as Type B response In rare cases where sensitive clays or structured soils are encountered a
Type C relationship occurs whereby the load-displacement curve reaches a peak followed by strain
softening In the one case observed in this study for Type C loading (Haga clay) the corresponding
ldquopeak capacityrdquo was reached at a pseudo-strain of sB = 4 as shown in Figure B16 followed by some
strain softening In the cases of soft clays at Belfast and Bothkennar a plunging type (plateau failure)
was achieved at around sB=7
B-20
Figure B15 Definitions of capacity from three types of observed foundation load test
responses (modified after Kulhawy 2004)
B-21
Figure B16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footings
In the case of fine-grained soils that develop excess porewater pressure during cone penetration the
use of the net total cone resistance (qtnet) which equals the cone tip resistance minus the total stress
must be considered (Lunne et al 1997) As noted earlier the correction of CPT data in sands is minimal
because of the low value of penetration porewater pressures and the fact that total overburden stress is
small relative to the cone resistance especially for shallow soundings Thus the use of qtnet may be
substituted for qc into the trend of Figure B1
B16 Undrained and Drained Capacity
The footings on sands typically show drained response with no excess porewater pressures developed
during loading For the foundations situated on silts analyses by Larsson (1997 2001) concluded that
these too show drained conditions For shallow foundations on clays the entire range of drainage cases
are possible ranging from undrained to partially-drained to fully-drained depending upon the applied
rate of loading relative to the permeability of the geomaterial If sufficient data were available for each
of the clay sites then the recent evaluation of drainage condition could be assessed using normalized
velocity criteria (Chung et al 2006) However this was not possible with the current data set Instead
B-22
the test loadings of footings on clays have essentially been assumed to primarily occur under undrained
conditions following recommendations by Tand et al (1986) For this case a limiting bearing stress can
be calculated from bearing capacity analyses and related directly to the CPT readings
From classical bearing capacity calculations using limit plasticity solutions the ultimate foundation stress
is shown in Equation B5
= 119925119940 ∙ 119956119958 B5 119954119958119949119957
where Nc equals the bearing factor for constant volume (514 for strip and 614 for square or circular
plan) and su equals the undrained shear strength of the clay For the case of soft to firm clays of low to
medium sensitivity the operational strength can be related to the preconsolidation stress (Mesri 1975
Jamiolkowski et al 1985) as shown in Equation B6
prime 119956119958 = 120782 120784120784120648119953 B6
Finally the effective preconsolidation stress has been linked directly to the net cone resistance (Chen amp
Mayne 1996 Demers amp Leroueil 2002) as shown in Equation B7
prime 120648119953 = 120782 120785120785(119954119957 minus 120648119959119952) B7
Combining Equations (B5) (B6) and (B7) results in direct expressions for undrained bearing capacity on
intact clays from CPT results as shown in Equations B8 and B8-2
Strip footing 119954119958119949119957 = 120782 120785120789120785(119954119957 minus 120648119959119952) B8
B-23
Squarecircle 119902119906119897119905 = 0445(119902119905 minus 120590119907119900) B8-2
Equation (B8-2) is in excellent agreement with Tand et al (1986) database study for the case of intact
clays and footings with no embedment Their study also provided a lower relationship for foundations
situated on jointed clays specifically recommending that the bearing stress not exceed 030 qtnet
Review of the available data in Figure 3 shows this to be rather conservative and a more realistic value
may be taken at 040 qtnet for fissured clays
B161 Normalized Undrained Footing Response
The characteristic stress vs normalized displacement curves for footings on clays exhibiting undrained
conditions can be verified using a relatively recent case study on Beaumont clay by Stuedlein amp Holtz
(2010) The Baytown Texas site was investigated using an exploration program of soil borings
piezocone soundings and laboratory testing The field testing included a large square footing (B = 276
m) and two small circular plates (D = 076 m) The measured load-displacement responses for all three
foundations are presented in Figure B17 Of course the large footing carried considerably higher loads
than the two small plates on the same soil deposit Yet using the aforementioned characteristic stress
(q) vs square root of normalized displacement (sB) essentially gave a unique relationship for all 3
foundations as presented in Figure B18 In this case utilizing the form of Equation B2 the
characteristic coefficient rs is 177 MPa for this deltaic clay formation
B-24
Figure B17 Measured load vs displacement curves for 3 shallow foundations on clay at
Baytown Texas (Stuedlein amp Holtz 2010)
B-25
Figure B18 Characteristic stress vs square root of normalized displacement for all three
foundations at Baytown site (Stuedlein amp Holtz 2010)
B17 General Direct CPT Method
In a manner analogous to the normalization of applied foundation stresses shown in Figure B14 for
footings on sands a generalized direct CPT method for shallow foundations on different soils can be
made in Equation B9
120782120787 119956 119954 = 119945119956 ∙ 119954119957119951119942119957 ∙ (119913
) lt 119954119940119938119953119938119940119946119957119962 B9
where qcapacity is equal to the foundation bearing capacity of the ground and hs is equal to the empirical
fitting term that depends on soil type Specifically hs equals 058 for sands 112 for silts 147 for
fissured fine-grained soils and 270 for clays A summary of the normalized stress versus the normalized
displacement curves for these four soil types is presented in Figure B19 where the data fitting to obtain
the hs parameter on clays are limited to (sB) lt 004 corresponding to undrained loading The data on
B-26
silts and sands are considered fully drained whereas the fissured clay subset may be partially drained to
undrained The statistical measures for obtaining the fitted parameter hs are quite good as shown in
Figure B20 with the coefficient of determinations (r2) values of 0947 for sands 088 for silts 0935 for
fissured and 0925 for intact clays Additional statistical evaluations on the database are provided by
Uzielli amp Mayne (2011)
Figure B19 Normalized foundation stress vs square root of normalized displacements
B-27
Figure B20 Applied foundation stress vs CPT-calculated stress for 67 footings
The same data are shown on Figure B21 in a log-log plot to emphasize the early parts of the load-
displacement responses of these footings The best fit line from regression analyses shows excellent
statistical correlations As detailed earlier the undrained bearing capacity of square or circular footings
on clays can be taken as approximately 040 times qtnet For silts and sands that experience drained
loading with no excess porewater pressures being developed and no clear peak value for capacity the
(sB) =10 criterion can be used In this case Equation 7 can be used directly to obtain stress q equal to
qcapacity when (sB)05 is 0316
An alternate approach is presented in Figure B22 summarizing the direct CPT method for estimating the
load-displacement-capacity of shallow square footings on four soil types The maximum values of soil
bearing capacity (qmax) can be capped at stresses corresponding to a percentage of the average
measured qtnet determined over the depth interval from the foundation bearing elevation to 15middotB
below the foundation In such an approach the foundation bearing capacity can be taken as
(qcapacityqtnet) of 02 for sands 035 for silts 040 for fissured and 045 for intact clays Using the assigned
values of the hs parameter the corresponding cut-off for the general stress-normalized displacement
relationship given by Equation 7 can be established specifically giving corresponding values of the
B-28
normalized displacements which can also be considered as pseudo-strains equal to (sB) max of 4 for
clays 7 for fissured clays 10 for silts and 12 for sands
Figure B21 Applied footing stress vs CPT calculated stress on logarithmic scales
B171 CPT Soil Material Index Ic
The soil behavioral type can be assessed indirectly by CPT using a material index (Ic) as defined by
Robertson (2009a b) shown in Equation B10
119920119940 = radic[120785 120786120789 minus 119949119952119944119928119957119951]120784 + [120783 120784120784 + 119949119952119944119917119955]120784 B10
B-29
Figure B22 Summary graph and equations of direct CPT method for evaluating the vertical
where Qtn equals the stress-normalized cone tip resistance and Fr equals the normalized sleeve friction
calculated from the cone penetrometer readings as shown in Equations B11 and B12 respectably
(119954119957minus120648119959119952)120648119938119957119950 = 119951 B11 119928119957119951 prime (120648119959119952120648119938119957119950)
119943119956 119917119955() = 120783120782120782 ∙ B12 (119954119957minus120648119959119952)
where σatm equals a reference stress equal to atmospheric pressure (σatm = 1 atm asymp 1 bar asymp 100 kPa) In
the initial evaluation the exponent n is set to n = 1 to find the soil behavioral type (SBT) based on a 9-
zonal chart as shown in Figure 23 The value of exponent n varies with soil type ranging from n = 1 in
intact clays and decreasing with increasing grain size to around approximately 075 in silts and
B-30
approximately 05 plusmn 02 in clean sands The appropriate value of exponent n is found by iteration until
conversion using the relationship (Robertson 2009b) shown in Equation B13
prime 120648119959119952 119951 = 120782 120785120790120783 ∙ 119920119940 + 120782 120782120787 ( ) minus 120782 120783120787 le 120783 120782 B13 120648119938119957119950
As the distinction that footings on intact clays subjected to fast loading are behaving primarily under
undrained conditions (ie constant volume) Robertson (2009a) suggested that this occurs when Ic gt 27
while in contrast Ic lt 25 corresponds more or less to drained loading (ie no excess porewater
pressures)
The CPT material index can be used to identify soil type and the corresponding characteristic hs value for
estimating footing load-displacement response In a number of the case studies investigated the results
of the sleeve friction readings were also available to allow the calculation of Ic with depth at these sites
Figure B23 shows the tentative relationship between the values of the hs parameter plotted versus the
corresponding CPT material index with an approximate trend given by Equation B14
120784120785 119945119956 = 120784 120790 minus 120783120787 B14
119920119940 120783+( ) 120784120786
B-31
Figure B23 Foundation soil formation parameter hs versus CPT material index Ic
Soil behavioral type from the Ic ranges given by Robertson (2009b) and are shown in Figure B23 with an
overall general agreement with the known soil classifications In summary the CPT can provide an
evaluation of the load-displacement-capacity response directly via Equation B15 which may be of
interest in mixed soil types such as silty sands sandy silts and the like
120784120785 Footing stress 119954 = 119954119951119942119957 ∙ radic(119956frasl119913) ∙ [120784 120790 minus
)120783120787] B15
120783+(119920119940frasl120784120786
As discussed earlier foundation bearing capacity may be taken by a limiting value of pseudo-strain (sB)
max or by qmax as specified as a percentage of qtnet This definition of bearing capacity can be seen in
comparison to soil type as seen in Figure B24
B-32
Figure B24 Bearing capacity ratio qmaxqnet versus CPT material index Ic
B172 Rectangular Foundations
This same elastic solution from Giroud (1968) for a CPT method for square and circular foundations were
utilized for rectangular foundations The study covered rectangular distortions (AB) ranging from 1
(square) to very long foundations with AB equal to 20 where A represents the foundation length and B
represents the foundation width (Mayne amp Dasenbrock 2017) The influence factor for rectangular
shaped foundations is given in Figure B25 and the expression in Equation B16
= (119912119913)120782120785120786120787 119920119912119913 B16
B-33
Figure B25 Influence factor for rectangular foundations from elastic theory solution (Mayne
amp Dasenbrock 2017)
Including 32 full scale load tests on sands results from an additional 98 shallow foundations have been
reported in previous studies The foundations were primarily square or circular and among these include
large spread footings with measured settlements and bearing capacity Table B4 shows a summary of
the number of footings and footing sizes used in the study The range of length to width ratios varied
from 1 lt (AB) lt 23 with an average value of (AB) equal to 238 The embedment depth to footing
width ranged from 0 lt DeB lt 222 and averaged 046
Table B4 Sources for shallow foundation performace
Source of Data No of
Footings
Mean B
(m)
Max B
(m)
Min B
(m)
Mean A
(m)
Max A
(m)
Min A
(m)
Mayne et al (2012) 32 149 609 046 149 609 046
Lehane (2011) 8 041 027 060 041 060 027
Gifford et al (1987) 17 846 1588 527 1591 3596 701
Jeyapalan amp Boehm
(1986)
13 1022 2892 400 1671 3049 640
B-34
Schmertmann
(1970)
31 945 5608 061 1288 8668 061
Papadopoulos
(1992)
29 870 3600 100 1300 7290 100
Total 130 671 5608 027 1012 8668 027
The solution for shallow foundation settlements given back in Equation B1 combined with
the expression in Equation B16 takes on the form
119954∙119913∙119920119912119913∙(120783minus120642120784) 119956 = B17
119916119956
Combining the direct CPT equation developed from the 32 full scale load tests (Equation B15)
together with the influence factor for rectangular shaped foundations becomes
minus120782120785120786120787 120784120785 119912 Footing stress 119954 = 119954119951119942119957 ∙ radic(119956frasl119913) ∙ [120784 120790 minus
)120783120787] ∙ [ ] B18
120783+(119920119940frasl120784120786 119913
B18 Conclusions
A direct CPT method for square rectangular and circular shallow footings is developed using a database
of 166 full-scale field load tests where the minimum footing size of an equivalent square width of 05 m
is established to avoid scaling issues The use of load vs displacement is generalized by characteristic
stress vs normalized displacement (sB) and further simplified by a square root plotting technique that
captures the ground response in a single parameter that is related to the qtnet Data are grouped
according to four main soil categories sands silts fissured clays and intact clays It is generally believed
that the footings on sands and silts exhibit fully drained behavior while in intact clays an undrained
response occurs under conditions of constant volume
B-35
The summary equation for evaluating the vertical stress-displacement-capacity of square rectangular
and circular footings is given by
minus120782120785120786120787 119956 119912 119954119943119952119952119957119946119951119944 = 119945119956 ∙ 119954119957119951119942119957 ∙ radic ∙ ( ) lt 119954119950119938119961 B19
119913 119913
where the parameter hs also relates directly to Ic The foundation capacity depends upon the mode of
failure as well as drainage conditions and can be taken as a fraction of qtnet where qmaxqtnet is 020 in
sands 035 in silts 040 in fissured clays and 045 in intact clays as shown in Figure 24 Alternatively the
capacity can be defined using a limiting pseudo-strain given by (sB) max of 4 in clays 7 in fissured
10 in silts and 12 in sands
B-36
APPENDIX C
List of Figures
Figure C1 Components of Axial Pile Capacity C-3
Figure C2 Selected alpha relationships as function of the undrained shear strengthC-5
Figure C3 Pile side friction coefficient β in terms of effective friction angle and overconsolidation ratio
C-11
Figure C4 Concept of Direct versus Rational Method for CPT evaluation of axial pile capacityC-14
Figure C5 Soil behavior type using CPT via original UniConeC-19
Figure C6 Soil behavior type using CPT via Modified UniCone C-19
Figure C7 Nine-zone soil behavioral soil type using normalized piezocone parameters C-21
Figure C8 Summary of Modified UniCone Method for direct CPT assessment of axial pile capacity C-22
Figure C9 Elastic continuum solution for axial pile response under compression and tension loadingC-24
List of Tables
Table C1 Alpha Methods for Axial Pile Side Friction where fp = αmiddotsu C-5
Table C2 Beta Methods for Axial Pile Side Friction where fp = βmiddotσvoC-9
Table C3 Selection of Direct CPT Methods for Axial Pile CapacityC-15
C-1
C1 EVALUATING DEEP FOUNDATION RESPONSE FROM CONE
PENETRATION TESTS
C11 Introduction
The axial response of deep foundations includes the load-displacement-capacity and axial load transfer
when driven pilings and drilled shafts are loaded in compression and uplift The use of cone penetration
testing (CPT) especially seismic piezocone tests (SCPTu) are advantageous since they provide
information on the subsurface soils and their geomaterial properties The SCPTu provides at least four
measurements on soil behavior with depth including (a) cone tip resistance qt (b) sleeve friction fs (c)
penetration porewater pressure u2 and (d) shear wave velocity Vs This offers the opportunity to
determine the geostratigraphy unit weight effective overburden stress shear strength stress state
and stiffness of the ground from a single exploratory sounding thus values in economic expediency
and reliability The axial pile capacity can be calculated based on static equilibrium of forces acting along
the sides and base of the pile foundation The displacement of the pile can be ascertained using
elasticity theory from closed-form solutions boundary elements andor finite element analyses Load
transfer occurs along the length of the pile and only a portion of axial forces are transmitted to the base
or toe or tip of the pile These too can be assessed using elasticity solutions
An alternate approach to pile capacity involves the utilization of direct CPT methods available
since the 1970s but in the past 10+ years a number of new and statistically reliable algorithms
have been developed which can be implemented for highway design and construction
C111 Axial Pile Capacity
The axial compression capacity of a single pile foundation is composed of a shaft or side component and
end-bearing component at the base as depicted in Figure 1 For a circular pile the side capacity (Qs) is
determined from the unit side friction (fp) acting along the surface area of the shaft which is As = πmiddotdmiddotL
where d = pile diameter and L = length embedded below grade
If the magnitude of side friction is uniform and constant with depth the side capacity is simply
C-2
119928119956 = 119943119953 middot 119912119956 C1
Moreover however many piles extend through multiple layers and a summation of unit side
frictions acting on various pile segments must be tabulated over the length of the pile as
suggested by Figure C1
Figure C1 Components of Axial Pile Capacity
The unit-end bearing resistance (qb) acts over the base of the pile tip where the area of a circular pile is
given by Ab = πmiddotd24 For piles in compression loading the base capacity is determined from Equation
C2
119876119887 = 119902119887 middot 119860119887 C2
and for piles in tension (or uplift) it is normally taken that Qb = 0
C-3
C112 Pile Unit Side Friction
Several different approaches can be adopted for evaluating the unit pile side friction (fp) prior to full-
scale load testing and construction (Poulos amp Davis 1980 ONeill 2001) The most common types
including the alpha and beta methods as summarized in Tables 1 and 2 respectively In the alpha
method an empirical coefficient (α) is applied to the undrained shear strength (su) of clay soils to
obtain
119891119901 = 120572 middot 119904119906 C3
An illustrative example of alpha curves is shown in Figure C2 A difficulty with the alpha
method is the evolution of many variants and changes to the expressions for its estimation It is
also restricted in its use for specific types of piles in clays and fine-grained soils
C-4
Figure C2 Selected alpha relationships as function of the undrained shear strength
Table C1 Alpha Methods for Axial Pile Side Friction where fp = αmiddotsu
C-5
Reference
Source
Alpha Equation Remarks
Tomlinson
(1957)
2716801102
uu ssAll pile types (steel
concrete timber)
Note su in ksf
Tomlinson
(1957)
2616201102
uu ssConcrete piles
Note su in ksf
McClelland
(1974)
Kerisel α = 07 - 031middotln(su)
Peck α = 10 - 02middotsu
Woodward α = 091middot(su)091
Driven piles in clay
Note su in ksf
Semple
(1980) )1(511
)1(1
OCR
OCR Summary from 9 series of
pile load tests
American
Petroleum
Institute (API
1981)
For suσvo le 035 α = 10
For suσvo gt 080 α = 05
Otherwise α = 1 + 1111middot (035 - suσvo)
Driven steel pipe piles
Tomlinson
(1986)
1 α = 55 (su)-091
2 α = 785 (su)-102
3 α = 605 (su)-100
where 75 lt su (kPa) lt 210
1 Driven piles
2 Bored piles
3 Driven piles in till with L
lt 10middotd
API (1987) 1 α = 10 for su lt 25 kPa Driven piles in clay other
than Gulf of Mexico
C-6
2 α = 125 - 001middotsu for 25 lt su lt 75 kPa
3 α = 050 for su gt 75 kPa
Kulhawy amp
Jackson
(1989)
α = 021+026middot(σatm su) le 1 Analyses of 106 drilled
shaft foundations
Table C1 Continued
API (1989) 05 For suσ vo le 1 α =
su radic frasl prime σvo
minus025 su For suσ vo gt 1 α = 05 ∙ ( frasl prime ) σvo
Driven steel pipe piles
Kulhawy amp
Jackson
(1989)
OCR
OCR
sin50
tan)sin1( sin
Λ = (1-CsCc) asymp 080 for
insensitive clays
Nowacki et al
(1996)
minus05 su For suσvo le 07 α = 05 ∙ ( frasl prime ) σvo
minus02 su For suσvo gt 07 α = 055 ∙ ( frasl prime ) σvo
Driven pile foundations
Kolk and van
der Velde
(1996)
α =09middot FL middot (suσvo)-03 lt 10
FL = [ (L - z)d]-02 = length term
L = pile length
Note su obtained from
UU lab tests
Applicable to driven piles
in clays
C-7
Knappett amp α = 1 for su le 30 kPa Non-displacement piles in
Craig (2012) fine-grained soils α = 116 - (su185) for 30 kPa le su le 150 kPa
α = 035 for su gt 150 kPa
Knappett amp
Craig (2012) α = 055 ∙ 02 40
( ) (LfraslD)
∙ su ( frasl prime σvo
minus03
) Displacement piles in fine-
grained soils
z = depth at point considered
d = outside diameter of pile
Karlsrud et al
(2005 NGI
Method)
α = function (suσvo and plasticity index) Driven pilings
Fleming et al
(2009) 05 minus05 su su For su σvo le 1 120572 = ( frasl prime ) ∙ ( frasl prime )
σvo NC σvo
05 minus025 su su For su σvo gt 1 α = ( frasl prime ) ∙ ( frasl prime ) σvo NC σvo
For driven piles in clay
where (suσvo)NC is the
normalized shear strength
for normally-consolidated
clay
Brown et al
(2010)
α = 0 for 0 lt z le 5 feet
α = 055 for z gt 5 feet and (suσatm) le 15
α = 055-01middot(suσatm -15) for 15le (suσatm) le 25
Drilled Shafts and Bored
Piles
Table C1 Continued
C-8
OCRK )sin1(0
Karlsrud
(2012)
α = function (su σvo and plasticity index) Driven pilings
Note as reported by Karlsrud (2012)
In the beta method the coefficient β is applied to the effective overburden stress (σvo) at the point of
concern along the pile length
prime = 120573 C4 119891119901 middot 120590119907119900
Table C2 Beta Methods for Axial Pile Side Friction where fp = βmiddotσvo
Reference Source Beta Equation Remarks
Burland (1973) β = (1-sinϕ) middot tan ϕ Piles in NC clays
Meyerhof (1976) β = Kmiddottanδ where K = K0 = for NC clays
K = 15middotK0 for OC clays
Poulos amp Davis
(1980)
β = (1-sinϕ) middot tan ϕ middot OCR05 Piles in OC stiff clays
Table C2 Continued
C-9
Kulhawy amp Jackson
(1989)
β = (1-sinϕ) middot tan ϕ middot OCRsinϕ Piles in quartz sands and insensitive
clays
ONeill (2001) β = (1-sinϕ) middot tan θ middot OCRsinϕ Drilled shafts where θ = (δϕ)middotϕ and
δ = interface friction between pile and soil
Fleming et al
(2009)
β = Kmiddottanδ K = lateral stress coefficient and δ =
friction angle between soil and pile
material
Karlsrud (2012) β = fctn (OCR and PI) PI = plasticity index of the clay ()
Mayne and Niazi
(2017)
β = CM middot CK middot K0 middot tanϕ
where K0 = (1-sinϕ) middot OCRsinϕ for soils
that are virgin loaded then unloaded
CM = pile material factor = 1 (drilled
augered) 09 (prestressed or precast
concrete) 08 (timber) and 07
(steel) and CK = pile installation factor
= 09 (bored or augered) 10 (low
displacement eg H-pile or open-end
pipe) and 11 (driven high
displacement eg prestressed
concrete closed-end pipe)
As the beta approach applies to all types of soils (gravels sands silts and clays) and more or
less has remained unchanged since its advent circa 1970 it has been selected for further
discussion herein In the direct form for calculation of unit pile side friction the expression is
given by
119874119862119877119904119894119899120601 prime prime 119891119901 = 119862119872 ∙ 119862119870 ∙ (1 minus 119904119894119899120601prime) middot ∙ 119905119886119899120601prime ∙ 120590119907119900 C5
C-10
where CM is a soil-pile interface friction coefficient and CK = pile installation factor Values of CM are in
the range 07 lt CM lt 10 depending upon pile material (steel wood concrete) and ranges for CK vary
09 lt CK lt 11 depending upon method of installation (auger drill driven) as detailed in Table 2
The value of K0 is limited to the passive stress coefficient which for the simple Rankine case is given by
KP = (1+sin ϕ) (1 - sin ϕ ) Thus there is a maximum value of overconsolidation ratio for which
Equation C5 applies given by OCRlimit = [(1+sin ϕ ) (1 - sin ϕ )2] (1sinϕ)
The corresponding graph in Figure C3 shows the relationship for β in terms of ϕ and OCR
Figure C3 Pile side friction coefficient β in terms of effective friction angle and overconsolidation ratio
C113 Pile Unit End Bearing
C1131 Theoretical Considerations
The evaluation of the end bearing resistance of pile foundations is commonly determined using
limit plasticity theory where
prime Drained loading = C6 119902119887 119873119902 middot 120590119907119900
C-11
Undrained loading 119902119887 = 119873119888 middot 119904119906 C7
where the value of σvo is calculated at the depth z = L and the value of su is the average undrained shear
strength from z = L to a depth z = L + d beneath the pile tip For a circular pile the limit plasticity solution
for undrained loading gives (Vesic 1977) a value Nc = 933 while a deep strip foundation would employ a
value of Nc = 824 For drained loading the expression for Nq for a deep foundation is given by (Vesic
1977)
)]arctan()sin1(tan21[)](tan1[sin1
sin1)tanexp( 2 BLABNq
C8
where A and B are the pile plan dimensions (for a circular pile A = B) and L = pile length For a
circular pile this can be approximated by
asymp 077(120601prime75deg) 119873119902 C9
over a range of effective friction angles 20deg le ϕ le 45deg
C1132 Practical Considerations
For drained loading of sands the full calculated end bearing capacity will never be realized because it
would require the pile to move a distance equal to its diameter This is beyond practical use and
therefore the limit plasticity solutions must be clipped to a fraction of the calculated value Another
reason for using a reduced end-bearing capacity is due to strain incompatibility since the side capacity is
mobilized early while the end-bearing is engaged much later So to have compatible values of Qs and
Qb the qb must be reduced using the following guidelines (Randolph 2003 Mayne 2007)
prime prime C10 119902119887 = 119873119902 ∙ 120590119907119900 ∙ 119891119909
where fx = strain incompatibility factor
fx = 010 for drilled shafts augered cast-in-place and bored piles
fx = 020 for driven low displacement piles (opened-ended steel pipe and H-piles)
fx = 030 for driven high-displacement piles (ie solid piles PSC and closed-ended pipe)
C-12
For undrained loading of circular piles in compression no reduction of end-bearing is necessary
therefore
119902119887 = 933 middot 119904119906 C11
C12 Direct CPT Methods for Pile Capacity
In the direct CPT method the penetrometer readings are scaled directly via specified algorithms to
obtain the pile unit side friction and end-bearing as depicted in Figure C4 As many as 40 different
direct CPT methods have been developed over the past five decades as summarized by Niazi amp Mayne
(2013) Starting circa 1970 many of these early methods relied on hand-recorded data from field
mechanical CPTs where only qc data were obtained at 20 cm intervals or later with mechanical readings
of both qc and fs using special sets of inner and outer rods that recorded vertical load for tip and loads
for tip plus sleeve in alternating increments Also early pile load test data were often obtained from
top-down measurements of load-displacement
C-13
Undrained qb = Nc su
Qside = S (fp DAs)
QTotal = Qs + Qb - Wp
fp = cmck Ko svorsquo tanrsquo
Qbase = qb Ab
Drained qb = Nq svorsquo
Method Two Rational or ldquoIndirectrdquo Method
Method OneldquoDirectrdquo CPT Method
(Scaled Pile)
fp = fctn (soil type pile
type qt fs and u2)
qb = fctn (soil type qt-u2)
qb = unit end bearing
unit sidefriction fp
OCR su Ko t DR rsquo
AXIAL PILE CAPACITY FROM CPT READINGS
Figure C4 Concept of Direct versus Rational Method for CPT evaluation of axial pile capacity
Beginning in the mid-1990s as the modern electric piezocone (CPTu) was implemented the newer
equipment offered improved data and better resolution because of the additional reading of porewater
pressures and correction of raw measured qc to total resistance qt In addition the use of electronic
digital data collection and field computers proved superior in field measurements and recordings As a
consequence several reliable CPT methods for axial pile capacity have been developed Moreover
parallel improvements in full-scale pile load testing occurred and now include modern strain gage
instrumentation digital data recording and automated testing procedures Also the testing can be
single direction (compression or tension) or bi-directional as in the Osterberg cell
C-14
Table C3 provides a selection of recent direct CPT methods that have become available over the
past two decades A number of these (namely ICP NGI UWA and Fugro) were funded by the
offshore industry because of the growth of oil amp gas reserves and windfarm installations thus
necessitating increased concerns on risk probability and reliability in the site investigations for
offshore platforms and design of large driven monopile foundations These direct CPT methods
are often statistically based on large datasets compiled from full-scale load tests made
worldwide (eg Schneider et al 2008)
Table C3 Selection of Direct CPT Methods for Axial Pile Capacity
Method Pile Types Soil Types References CPT
data
Additional
parameters
needed
Unicone Method all types Sands silts
clays
Eslami and
Fellenius
(1997)
Fellenius
(2009)
qt fs u2
KTRI = Kajima
Technical
Research
Institute
all types Sands
mixed clays
Takesue et al
(1998)
fs and u2 No guidance given
on end bearing
resistance
Imperial College
Procedure (ICP)
OE and CE Sands Chow et al
(1997 PhD)
Jardine et al
(2005)
qt Interface friction
angle (δ) from
ring shear tests
Correlated to
mean grain size
(D50)
C-15
OE and CE Clays Jardine et al
(2005)
qt Interface friction
angle (δ)
correlated to PI
Table C3 Continued
NGI Method
(Norwegian
Geotechnical
Institute)
OE and CE Sands Clausen et al
(2005)
qt DR from qt
Clays a Alpha
method
(Karlsrud et al
2005 2012)
b Beta
method
(Karlsrud
2012)
qt for su
qt for
OCR
PI = plasticity
index ()
PI = plasticity
index ()
Fugro Method OE and CE Sands Kolk et al
(2005)
qt
Clays Van Dijk and
Kolk (2011)
qt
OE and CE Sands Lehane et al
(2005)
qt IFR = infilling ratio
related to amount
of plugging
C-16
Purdue LFRD Drilled
shafts
Sands Basu amp
Salgado
(2012)
qt LRFD = load
resistance
factored design
Enhanced Various Sands silts Niazi and qt u2 Based on 330 load
Unicone Method clays and Mayne (2015 and fs tests including
mixed soils 2016) bored augered
jacked and driven
piles
UWA (Univ of
Western
Ra = pile
roughness
Australia) Clays Lehane et al
(2012)
qt Interface friction
angle (δ) from
ring shear tests
and also method
without δ
HKU Method
(Hong Kong
Univ)
OE and CE
(end
resistance
only)
Sands Yu and Yang
(2012)
qt End bearing only
PLR = plug length
ratio
Note PLR can be
estimated from
OE inner diam
Table C3 Continued
Note previously called Marine Technical Directorate (MTD)
C-17
Many of the offshore CPT methods relate to driven piles either in sand or clay as that is
common deep foundation for those purposes Of particular interest are the Unicone and
Modified Unicone Methods since they use all three readings of the piezocone (qt fs and u2)
and address a variety of pile foundation types
C13 Modified UniCone Method
The modified UniCone Method was developed based on a total 330 pile load tests which were
associated with SCPTu data during their site investigations (Niazi and Mayne 2015 2016) This
represents a threefold increase over the original UniCone database that was built upon data
from 106 pile load tests (Eslami amp Fellenius 1997)
For the original UniCone algorithms use is made of the effective cone resistance (qE)
119902119864 = 119902119905 minus 1199062 C12
and a chart of qE vs fs provided an approximate soil classification in five distinct groups as shown by
Figure C5 Later in the modified approach a better delineation of the larger dataset gave soil
subclassifications as indicated by Figure C6
C-18
Figure C5 Soil behavior type using CPT via original UniCone
Figure C6 Soil behavior type using CPT via Modified UniCone
C-19
In the modified approach the 9-zone normalized soil behavioral type (SBTn) is ascertained using CPT
data in conjunction with the Robertson (2009) charts as shown in Figure C7 As discussed in Appendix
A and B the SBTn method uses the normalized cone resistance (Qtn) normalized sleeve friction (Fr())
and CPT material index Ic This permits a much wider range in the calculated pile side friction because fp
is related as a continuous curve with Ic rather than only 5 values that are assigned in the original
scheme
The pile unit side friction (fp) is obtained from qE and the CPT material index Ic using the following
expression at each elevation along the sides of the pile
10(0732 middot 119868119888 minus 3605) fp middot middot C13 = 119902119864 middot 120579119875119879 middot 120579119879119862 120579119877119860119879119864
C-20
Figure C7 Nine-zone soil behavioral soil type using normalized piezocone parameters
(after Robertson 2009 Mayne 2014)
where θPT = coefficient for pile type (θPT = 084 for bored piles 102 for jacked piles 113 for driven
piles) θTC = coefficient for loading direction (θTC = 111 for compression and 085 for tension) and θRATE
= rate coefficient applied to soils in SBT zones 1 through 7 (θRATE = 109 for constant rate of penetration
tests and 097 for maintained load tests) Since CPT provides data at regular intervals of 2 cm to 5 cm
along the sides of the pile the average fp from z = 0 to z = L can be used directly in Equation 1 to obtain
the shaft capacity Qs
C-21
The pile end bearing resistance is obtained from
middot 10(0325middot 119868119888minus1218) 119902119887 = 119902119864 C14
where qE is averaged in the vicinity of the pile tip Figure C8 shows the Modified Unicone Method in
graphical format For sensitive clays of zone 1 please see additional discussions by Niazi amp Mayne
(2016)
Figure C8 Summary of Modified UniCone Method for direct CPT assessment of axial pile
capacity
C-22
C14 Axial Pile Displacement
The movement of pile foundations can be assessed using elastic continuum theory (Poulos amp
Davis 1980 Randolph 2003 Mayne amp Niazi 2017) These relationships have been developed
using finite element analyses boundary elements and analytical closed-form solutions For the
latter approach the top displacement of a rigid pile subjected to a vertical force is shown in
Figure C9 This gives the movement at the top of a pile subjected to either compression or
tension (uplift) loading Also the percentage of axial load transferred to the pile toe is
determined For piles extending through various soil layers the elastic solution can be
implemented by using a set of stacked pile segments each with its own stiffness as
represented by a soil Youngs modulus
For pile groups the use of computer software is recommended Several available programs can handle
pile groups under axial and lateral moment loading such as DEFPIG (Univ Sydney) and PIGLET (Univ
Western Australia) A full listing of pile foundation software is given at the Geotechnical amp
GeoEnvironmental Service Directory wwwggsdcom
Ps
= Sh
aft
Load
sL
t
tEd
IPw
Load Transfer to Base
)]1)((5ln[
)(
)1(1
1
2 vdL
dLFI
E
ECT
211
IF
P
P CT
t
b
Base Load = Pb
Randolph Elastic Solution
Ground Surface
Top Load Pt = Ps + Pb
Rigid Pile Response
wt = displacementd = diameter
L = length
Es = Elastic Soil Modulus
Es0 (at z = 0) surface
EsM (at z = L2)mid-length
EsL (at z = L)full length
Load Transfer to Shaft
E = EsMEsL = Gibson parameterE = 1 for homogeneous case (constant E)E = 05 for pure Gibson case (Es0 = 0)
where FCT = load direction (= 1 compression 0 tension)
21
IF
P
P CT
t
b
z = depth
= Poissons ratio
Tension
Compression
C-23
Figure C9 Elastic continuum solution for axial pile response under compression and tension
loading
C-24
C15 Acknowledgements
The author appreciates the help of Fernando Illingworth of PonsEcuador Meena Viswanth of
GeoSyntecAtlanta and Dr Marco Uzielli of GeoRiskItaly in helping to collect and process the data
Funding support was provided by David Woeller of ConeTec Richmond BC
C-25
- Structure Bookmarks
-
- CONE PENETRATION TEST DESIGN GUIDE FOR STATE GEOTECHNICAL ENGINEERS
- TABLE OF CONTENTS
- LIST OF TABLES
- CHAPTER 1 INTRODUCTION
- CHAPTER 2 DIRECT CPT METHOD FOR SOIL CHARACTERIZATION
- CHAPTER 3 DIRECT CPT METHOD FOR SHALLOW FOUNDATIONS
- CHAPTER 4 DIRECT CPT METHOD FOR DEEP FOUNDATIONS
- REFERENCES
- APPENDIX A
- APPENDIX B
- APPENDIX C
-
313 Step 2a Foundation soil formation parameter59
314 Step 3 Soil elastic modulus and Poissonrsquos ratio 60
315 Step 4 Net cone tip resistance 60
316 Step 5 Bearing capacity of the soil 61
317 Step 6 Settlement61
318 Step 7 Final Check 61
32 Example Problems 62
321 Example 3 Direct CPT Method on Sands 62
CHAPTER 4 DIRECT CPT METHOD FOR DEEP FOUNDATIONS 70
41 Introduction70
42 Modified UniCone Method72
421 Step 172
422 Step 272
423 Step 373
424 Step 473
43 Axial Pile Displacements 73
44 Example Problems 74
441 Example 5 Direct CPT Methods Axial Pile Capacity74
REFERENCES 87
APPENDIX A
APPENDIX B
APPENDIX C
1
2
3
4
5
6
7
8
9
10
11
12
13
14
15
16
17
18
19
20
21
22
23
24
LIST OF FIGURES
Figure Geoparameters determined from CPT 1
Figure Conventional method for shallow foundation design compared to direct CPT method 2
Figure Direct CPT evaluation of axial pile capacity 3
Figure Soil unit weight from CPT sleeve friction 6
Figure Approximate ldquorules of thumbrdquo method using CPT sounding from Wakota Bridge MN 7
Figure SBT zones using CPT Ic 9
Figure Yield stress exponent compared to CPT material index 11
Figure k vs dissipation time for 50 consolidation (Mayne 2017) 15
Figure CPT data from Benton County Minnesota for example problem 1 16
Figure Soil layers using ldquorules of thumbrdquo17
Figure Soil layer 1 using SBT method 23
Figure Soil layer 2 using SBT method 24
Figure Soil layer 3 using SBT method 25
Figure Soil layer 4 using SBT method 26
Figure CPT data from Minnesota for example problem 2 34
Figure Soil layers using ldquorules of thumbrdquo36
Figure Soil layer 1 using SBT method 42
Figure Soil layer 2 using SBT method 43
Figure Soil layer 3 using SBT method 44
Figure Soil layer 4 using SBT method 45
Figure Dissipation t50 data 53
Figure Direct CPT method for shallow foundations56
Figure Direct CPT method introduction 58
Figure Differentiation of porewater pressure measurement locations (Lunne et al 1997) 58
Figure 25 Foundation soil formation parameter hs versus CPT material index Ic (Mayne 2017) 60
Figure 26 Diagram of footing profiles for Example 362
Figure 27 CPT data from Northern Minnesota 63
Figure 28 Schematic of foundation design associated with CPT data 64
Figure 29 Soil type for example problem 367
Figure 30 Direct CPT evaluation of axial pile capacity 70
Figure 31 UniCone Method soil behavior type using CPT (Mayne 2017) 71
Figure 32 Modified UniCone Method soil behavior type using CPT (Mayne 2017) 71
Figure 33 Deep foundation end bearing pile diagram74
Figure 34 CPT data from Minnesota for example problem 5 75
Figure 35 Soil layers using ldquorules of thumbrdquo for pile capacity example using direct CPT method 76
Figure 36 Soil layer 1 using SBT method 81
Figure 37 Soil layer 2 using SBT method 82
Figure 38 Soil layer 3 using SBT method 83
Figure 39 Soil layering compared to pilings 85
LIST OF TABLES
Table 1 Geoparameters calculated directly from CPT 4
Table 2 Minor Geoparameters 5
Table 3 Soil type compared to exponent m 10
Table 4 Geoparameters evaluated for Case Example 117
Table 5 Geoparameters 35
Table 6 Bearing capacity defined by soil type61
Table 7 SCPT Results 62
CHAPTER 1 INTRODUCTION
The purpose of this manual is to provide an analysis of soil characterization shallow footings and deep
foundations using direct cone penetration testing (CPT) methods Geotechnical site characterization is
important for evaluating soil parameters that will be used in the analysis and design of foundations
retaining walls embankments and situations involving slope stability Common practice is to determine
these soil parameters through conventional lab and in-situ testing An alternative method uses CPT
readings of cone tip resistance (qt) sleeve friction ( fs) and pore pressure (u2) directly to determine these
parameters such as unit weight effective friction angle undrained shear strength and many others
(Figure 1) Direct CPT methods are provided for these parameters which will be used in the designs for
shallow and deep foundations
Figure 1 Geoparameters determined from CPT
Designing shallow foundations is typically done in a two-part process determining the bearing capacity
and expected settlement (commonly referred to as displacement) of the soil to approximate the
required size and shape of a foundation The older traditional methods are no longer required with
many approaches existing for using CPT directly in the design of shallow foundations Results from these
methods can provide a direct assessment of bearing capacity andor settlement A specific approach to
the direct method has been recommended and tested using a database of 166 full-scale field load tests
(Figure 2) A one-part process is used to scale the measured cone penetrometer readings (ie
measured cone tip resistance sleeve friction and pore water pressure) to obtain the bearing capacity
and settlement of the soil A step-by-step procedure has been created to transition from the CPT data to
bearing capacity with settlement accounted for The steps consist of estimating a design footing width
1
and length while using the process in the Soil Characterization section to determine soil parameters
directly from the CPT data
Figure 2 Conventional method for shallow foundation design compared to direct CPT method
There are upwards of 40 different direct CPT methods that have been developed over the past five
decades to determine the axial compression capacity of a piling foundation Earlier direct CPT methods
relied on hand-recorded information where mechanical-type CPT cone tip resistance data would be
collected at 20 cm intervals
Herein the method recommended for deep foundation design is the Modified UniCone method which
uses all three readings of the modern electronic piezocone penetrometer (CPTu) while addressing a
variety of pile foundation types (Figure 3) The modified UniCone Method is based on a total of 330 pile
load tests (three times the original UniCone database) that were associated with SCPTu data Two
computer software programs have been recommended for analyzing pile movements andor
settlements
2
Figure 3 Direct CPT evaluation of axial pile capacity
3
Symbol Parameter
γt Soil total unit weight
Ic CPT material index
SBT Soil behavior type (SBT)
ʹ σp Preconsolidation stress
YSR Yield stress ratio
ϕʹ Effective friction angle
Ko Lateral stress coefficient
su Undrained shear strength
Dʹ Constrained modulus
Eʹ Drained Youngrsquos modulus
CHAPTER 2 DIRECT CPT METHOD FOR SOIL
CHARACTERIZATION
21 INTRODUCTION
A direct CPT method for determining the value of each of various geoparameters shown in Table 1 is
provided The parameter Ic is used in the derivation of several sequential parameters This section of
the guide may be referred to as these parameters are used in calculations throughout the shallow
foundations and deep foundations sections
Table 1 Geoparameters calculated directly from CPT
4
Kʹ Bulk modulus
ks Subgrade reaction modulus
MR Resilient modulus
Gmax Small-strain shear modulus
k Coefficient of permeability
cv Coefficient of consolidation
Symbol Parameter Equation
ρt Mass Density ρt = γtga
where ga = 98 ms2
σvo Total Stress σvo asymp Σ (γti middot Δzi)
c Effective cohesion ʹ Empirical c asymp 003σp
In clays c asymp 01cu
Other minor parameters can be used in calculations of the geoparameters in Table 1 These minor parameters are provided in Table 2
Table 2 Minor Geoparameters
5
22 SOIL UNIT WEIGHT
The total soil unit weight can be estimated from CPT sleeve friction resistance as shown in Figure 4
(Mayne 2014) This method is not applicable to organic clays diatomaceous soils peats or sensitive
soils
Figure 4 Soil unit weight from CPT sleeve friction
Use Equation 1 to calculate the soils total unit weight
1 120574119905 = 120574119908 ∙ [122 + 015 ∙ ln (100 ∙ 119891119904 + 001)]
120590119886119905119898
Since soil unit weight is required for determining most geoparameters (including Soil Behavior Type)
estimating the soil type of the layers using the ldquorules of thumbrdquo method is a good first step before
determining more precise layering (Figure 5) Once a unit weight is determined for each noticeable layer
change these results can be used in later calculations such as ldquoCPT Material Indexrdquo To use the ldquorules
of thumbrdquo method some helpful guidelines are to assume sands are identified when qt gt 725 psi and u2
asymp uo while the presence of intact clays are prevalent when qt lt 725 psi and u2 gt uo The magnitude of
porewater pressures help to indicate intact clays such as soft (u2 asymp 2middotuo) firm (u2 asymp 4middotuo) stiff (u2 asymp 8middotuo)
and hard (u2 asymp 20middotuo) Fissured overconsolidated clays tend to have negative u2 values such that u2 lt 0
6
Figure 5 Approximate ldquorules of thumbrdquo method using CPT sounding from Wakota Bridge MN
23 CPT MATERIAL INDEX
The development of the CPT material index Ic has improved the initial classification of soil types and
calculation of soil parameters as shown in Table 1 To calculate Ic follow steps 1 and 2
231 Step 1 Normalized Sleeve Friction
Calculate Fr using Equation 2 if not provided with the CPT data gathered
2
7
119891119904 119865119903() = 100 ∙
(119902119905minus120590119907119900)
232 Step 2 Iteration
Iterate using Equations 3-5 to determine Ic by initially using n = 1 to calculate a starting value of Ic The
exponent n is soil-type dependent n = 1 (clays) n asymp 075 (silts) and n asymp 05 (sands) Iteration converges
quickly which is generally after the 3rd cycle
3 (119954119957minus120648119959119952)120648119938119957119950 = 119951 119928119957119951 prime (120648119959119952120648119938119957119950)
4
prime 120590119907119900 119899 = 0381 ∙ 119868119888 + 005 ( ) minus 015 120590119886119905119898
119899 le 10
5 119920119940 = radic[120785 120786120789 minus 119949119952119944119928119957119951]120784 + [120783 120784120784 + 119949119952119944119917119955]120784
24 SOIL BEHAVIOR TYPE (SBT)
Typically soil samples are not taken when using CPT The soil types are then inferred from the qt fs and
u2 readings To determine the types of soil from CPT data follow steps 1 and 2
241 Step 1
To determine the soil layers from CPT results calculate Ic by following the steps under section ldquoCPT
Material Indexrdquo After Ic has been determined through all specified depths use Figure 6 to classify the
type of soil by comparing each Ic value to normalized CPT readings (Fr and Qtn) from Equations 2 and 3
8
Figure 6 SBT zones using CPT Ic
242 Step 2
To determine if any of the soil layers contain ldquosensitive clays and siltsrdquo from zone 1 or ldquovery stiff
overconsolidated (OC) soilrdquo from zones 8 and 9 use Equations 6 and 7 If any soil layers are found within
zone 1 by Equation 6 then caution should be taken as these clays are prone to instability collapses and
difficulties in construction performance Very stiff OC sands to clayey sands of zone 8 (15 lt Fr lt 45)
and very stiff OC clays to silts of zone 9 (Fr gt 45) can be identified by Equation 7
6 Equation 6 errata This is an exponential expression (see Figure 5)
119876119905119899 lt 12119890119909119901(minus14 ∙ 119865119903)
7 119876119905119899 gt 0005(119865119903minus1)minus00003(119865119903minus1)2minus0002
1
Errata See terms and coefficients in Figure 5 above (numbers cannot be rounded off)
After each CPT reading has been assigned a zone from Figure 14 a visual representation can be made to
show the predominant layers by soil types (Figure A5)
25 EFFECTIVE STRESS FRICTION ANGLE
The effective friction angle (ϕ) is used to govern the strength for sands and clays where Equation 8 is
used for sands and Equation 9 is used for clays The value of ϕ for sands is derived from Ic so before ϕ
can be calculated refer to the iteration of Qtn under the ldquoCPT Material Indexrdquo section Once the values
9
Soil Type m
Fissured clays 11
Intact clays 10
Sensitive clays 09
Silt mixtures 085
Silty sands 080
Clean sands 072
Note m may be higher than 11 in fissured clays
of Ic and Qtn are calculated the type of soil can be determined from the section ldquoSoil Behavior Type
SBTrdquo The type of soil will dictate which equation to use for ϕ
Sands
8 120601prime(deg) = 176deg + 110deg log(119876119905119899)
Clays
9 119902119905minus120590119907119900 120601prime(deg) = 295deg ∙ 119861119902
0121 ∙ [0256 + 0336 ∙ 119861119902 + log ( )] prime 120590119907119900
Where = 119861119902 (1199062 minus 119906119900)(119902119905 minus 120590119907119900)
26 STRESS HISTORY
Determining the stress history can be characterized by an apparent yield stress ratio of the form
10 prime 120590119901
119884119878119877 = prime 120590119907119900
Where σp is defined as the preconsolidation stress or effective yield stress (Equation 10) The YSR is the
same equation as the more common overconsolidation ratio (OCR) but is now generalized to
accommodate mechanisms of preconsolidation such as ageing desiccation repeated cycles of wetting-
drying repeated freeze thaw cycles and other factors
11 prime prime 120590119901 = 120590119910 = 033(119902119905 minus 120590119907119900)119898prime
Where σp σy σvo and qt have units of kPa and the value of m depends on soil type with typical values shown in Table 3
Table 3 Soil type compared to exponent m
10
The value of m for non-fissured soils and inorganic clays and silts is derived from Ic (Figure 7) so before
m can be calculated (Equation 12) refer to the iteration of Ic under the ldquoCPT Material Indexrdquo section
Determine Ic for all soil layers before calculating m
12 028
119898prime = 1 minus 25 119868119888 1+( frasl ) 265
Once m is known for all soil layers the YSR can be determined
Figure 7 Yield stress exponent compared to CPT material index
A limiting value of YSR can be reached for clays and sands It can be calculated using Equation 13
13 (1 sin(emptyprime))
(1+sin(emptyprime)) 119884119878119877119897119894119898119894119905 = [ ]
(1minussin(emptyprime))2
27 LATERAL STRESS COEFFICIENT
The lateral stress coefficient Ko = σhoσvo commonly referred to as the at-rest condition is used to
represent the horizontal geostatic state of soil stress Ko can be calculated using Equation 14 but
11
14
119870119900 = (1 minus sin(emptyprime)) ∙ 119884119878119877sin( emptyprime)
sections such as ldquoStress Historyrdquo and ldquoEffective Stress Friction Anglerdquo will need to be referred to for
determining parameters ϕ and YSR
A maximum value for Ko can be determined by Equation 15
15 (1+sin(emptyprime))
119870119900119898119886119909 = = 1199051198861198992(45deg + emptyprime2) (1minussin(emptyprime))
28 UNDRAINED SHEAR STRENGTH
Loading on soils can result in fully drained partially drained or fully undrained conditions Sands
typically produce drained cases due to their high permeability but exceptions may occur in loose sands
during fast loading where the water does not have sufficient time to dissipate Clays exhibit low
permeability and thus often result in undrained loading cases when a load is applied quickly For soft-
firm clays the undrained shear strength (su) can be determined from CPT via Equation 16 where the
value of the bearing factor Nkt can be taken as 12
16 119902119905minus120590119907119900 119904119906 =
119873119896119905
In the case of remolded undrained shear strength from CPT su asymp fs
29 GROUND STIFFNESS AND SOIL MODULI
Determining the grounds stiffness can be measured from geoparameters such as the constrained
modulus (D) drained Youngrsquos modulus (E) bulk modulus (K) subgrade reaction modulus (ks) resilient
modulus (MR) and small-strain shear modulus (Gmax)
17 119863 prime asymp 5 ∙ (119902119905 minus 120590119907119900)
18
12
119864 prime = 119863 prime
11
19 119864prime
119870prime = [3∙(1minus2119907prime)]
The subgrade modulus (ks) is a combination of soil-structural properties which creates a parameter that
depends on the ground stiffness and the size of the loaded element
20 119864prime
119896119904 = [119889∙(1minus1199072)]
The resilient modulus MR applies to pavement analysis and design and can be calculated using Equation
21 where qt and fs are in MPa
21 053 119872119877 = (146119902119905 + 135511989111990414 + 236)244
The small strain shear modulus Gmax is a representation of the initial stiffness of all soils and rocks
Graphically it is the beginning portion of all stress-strain-strength curves for geomaterials Use Equation
22 to determine Gmax
22 119866119898119886119909 = 120588119905 ∙ 1198811199042
Where Vs = shear wave velocity (Equation 23) as measured by seismic cone penetration tests (SCPT) If
only cone penetration tests (CPT) or piezocone (CPTu) data are available the shear wave velocity may
be estimated from
23 03 119891119904 119881119904 (119898frasl119904) = [101 ∙ log(119902119905) minus 114]167 ∙ (100 ∙ )
119902119905
Where qt and fs have units of (kPa)
210 COEFFICIENT OF CONSOLIDATION
The coefficient of consolidation (cv) controls the rate that foundation and embankment settlements
occur By using results of CPT dissipation tests that measure the change in u2 readings over time the
value of cv can be determined Using Equation 24 cv can be determined based on piezocone dissipation
13
curves The equation below requires an estimate of the in-situ rigidity index (IR) of the soil If results of
SCPTU are available then IR may be determined from Gmax and qt per equation A38
24 0030∙(119886119888)2∙(119868119877)075
119888119907 = 11990550
Where ac = penetrometer radius (178 cm for 10-cm2 cone 220 cm for 15-cm2 cone) t50 = time to reach 50 dissipation IR = Gsu = undrained rigidity index G can be determined from the ldquoGround Stiffness and Soil Modulirdquo section
211 HYDRAULIC CONDUCTIVITY
The hydraulic conductivity also known as the coefficient of permeability (k) expresses the flow
characteristics of soils and has units of cms or feetday One method of calculation would be to use
Equation 25 where cv would need to be determined from ldquoCoefficient of Consolidationrdquo and D would
need to be determined from ldquoGround Stiffness and Soil Modulirdquo
25 119888119907∙120574119908 119896 =
119863prime
An alternative approach developed for soft normally-consolidated soils (Figure 8) is shown in Equation
26 where t50 (sec) values are used directly in assessing k in (cms)
26
14
125
119896 asymp ( 1
) 251∙11990550
Figure 8 k vs dissipation time for 50 consolidation (Mayne 2017)
212 EXAMPLE PROBLEMS
2121 Example 1 Direct CPT Methods for Geoparameters on Sands
Several geoparameters need to be determined based on the given CPT data collected for the South
Abutment of a bridge in Benton County MN (Figure 9) The groundwater table (GWT) was measured at
17 feet Determine all the geoparameters found in Table 4 at depths of 0 feet to 30 feet All sand layers
can be assumed ldquodrainedrdquo with ν = 02
15
Figure 9 CPT data from Benton County Minnesota for example problem 1
16
Table 4 Geoparameters evaluated for Case Example 1
Symbol Parameter Symbol Parameter
γt Soil total unit weight Ko Lateral stress coefficient
Ic CPT material index Dʹ Constrained modulus
SBT Soil behavior type (SBT) Eʹ Drained Youngrsquos modulus ʹ σp Preconsolidation stress Kʹ Bulk modulus
YSR Yield stress ratio MR Resilient modulus
ϕʹ Effective friction angle Gmax Small-strain shear modulus
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 10 Soil layers using ldquorules of thumbrdquo
γw = 6224 pcf
17
Layer 1
From Figure 10 fs = 17 psi taken as a representative value of the layer
17 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1204 pcf
145 psi
Layer 2
From Figure 10 fs = 17 psi taken as a representative value of the layer
17psi
γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1204 pcf 145 psi
Layer 3
From Figure 10 fs = 12 psi taken as a representative value of the layer
12 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 4
From Figure 10 fs = 7 psi taken as a representative value of the layer
CPT Material Index
7 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1121 pcf
145 psi
Layer 1
From Figure 10 fs = 17 psi and qc = 3500 psi
18
qt = qc + u2 ∙ (1 minus a) = 3500 psi + 3 psi ∙ (1 minus 08) = 3501 psi
= γt ∙ 6 feet = 1204 pcf ∙ 6 feet = 7225 psf = 50 psi σvo
fs 17 psi Fr() = 100 ∙ = 100 ∙ = 049
(qt minus σvo) (3501 psi minus 50 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 5 psi minus 0 psi = 50 psi
(qt minus σvo)σatm (3501 psi minus 50 psi )145 psi = = Qtn prime )n (σvoσatm (50 psi145 psi)n
prime σvo 50 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(049)]2
Qtn = 35329 n = 036 lt 10 Ic = 13
Layer 2
From Figure 10 fs = 17 psi and qc = 3500 psi
19
qt = qc + u2 ∙ (1 minus a) = 3500 psi + 0 psi ∙ (1 minus 08) = 3500 psi
= 7225 psf + γt ∙ 6 feet = 7225 psf + 1204 pcf ∙ 6 feet = 1445 psf = 100 psi σvo
fs 17 psi Fr() = 100 ∙ = 100 ∙ = 049
(qt minus σvo) (3500 psi minus 100 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 100 psi minus 0 psi = 100 psi
(qt minus σvo)σatm (3500 psi minus 100 psi )145 psi = = Qtn prime )n (σvoσatm (100 psi145 psi)n
prime σvo 100 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
47 minus log(Qtn)]2 + [122 + log(049)]2
Ic = radic[3
Qtn = 27947 n = 041 lt 10 Ic = 14
Layer 3
From Figure 10 fs = 12 psi and qc = 1500 psi
20
qt = qc + u2 ∙ (1 minus a) = 1500 psi + 0 psi ∙ (1 minus 08) = 1500 psi
= 1445 psf + γt ∙ 11 feet = 1445 psf + 1172 pcf ∙ 11 feet = 2734 psf = 190 psi σvo
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 081
(qt minus σvo) (1500 psi minus 190 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = γwater ∙ (z minus 119911119908) = 6224 pcf ∙ (23 feet minus 17 feet) = 3734 psf = 99 psi
prime σvo = σvo minus uo = 190 psi minus 99 psi = 164 psi
(qt minus σvo)σatm (1500 psi minus 190 psi )145 psi = = Qtn prime )n (σvoσatm (164 psi145 psi)n
prime σvo 164 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(081)]2
Qtn = 947 n = 06 lt 10 Ic = 19
21
Layer 4
From Figure 10 fs = 7 psi and qc = 1200 psi
qt = qc + u2 ∙ (1 minus a) = 1200 psi + 0 psi ∙ (1 minus 08) = 1200 psi
= 2734 psf + γt ∙ 11 feet = 2734 psf + 1121 pcf ∙ 7 feet = 3519 psf = 244 psi σvo
fs 7 psi Fr() = 100 ∙ = 100 ∙ = 060
(qt minus σvo) (1200 psi minus 244 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = γwater ∙ (z minus 119911119908) = 6224 pcf ∙ (30 feet minus 17 feet) = 8091 psf = 130 psi
prime σvo = σvo minus uo = 244 psi minus 130 psi = 188 psi
(qt minus σvo)σatm (1200 psi minus 244 psi )145 psi = = Qtn prime )n (σvoσatm (188 psi145 psi)n
prime σvo 188 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(060)]2
22
Qtn = 686 n = 064 lt 10 Ic = 19
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic Qtn and Fr the first layer is defined as a ldquoDrained Gravelly Sandrdquo from
Figure 11
Figure 11 Soil layer 1 using SBT method
23
Layer 2
Based on values of Ic Qtn and Fr the second layer is defined as a ldquoDrained Sandrdquo from Figure
12
Figure 12 Soil layer 2 using SBT method
24
Layer 3
Based on values of Ic Qtn and Fr the third layer is defined as a ldquoDrained Sandrdquo from Figure 13
Figure 13 Soil layer 3 using SBT method
25
Layer 4
Based on values of Ic Qtn and Fr the fourth layer is defined as a ldquoDrained Sandrdquo from Figure 14
Figure 14 Soil layer 4 using SBT method
Effective Stress Friction Angle
Layer 1
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(3533) = 456deg
Layer 2
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(2795) = 445deg
26
Layer 3
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(947) = 393deg
Layer 4
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(686) = 378deg
Stress History
Layer 1
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (13frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(24139 kPa minus 345 kPa)072 = 4717 kPa
= 684 psi
prime σp 684 psi YSR = = = 136
primeσvo 50 psi
Layer 2
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + (14 1 + ( frasl ) frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(34132 kPa minus 689 kPa)072 = 605 kPa
= 877 psi
prime σp 877 psi YSR = = = 87
primeσvo 100 psi
27
Layer 3
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + (19 1 + ( frasl ) frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(10342 kPa minus 131 kPa)072 = 254 kPa
= 368 psi
prime σp 368 psi YSR = = = 22
primeσvo 164 psi
Layer 4
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (19frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(8274 kPa minus 168 kPa)072 = 215 kPa = 312 psi
prime σp 312 psi YSR = = = 17
primeσvo 188 psi
Lateral Stress Coefficient
Layer 1
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(456deg)) ∙ 136sin( 456deg) = 18
Layer 2
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(445deg)) ∙ 87sin( 445deg) = 14
28
Dprime 17480 psi
Eprime = = = 15890 psi 11 11
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(393deg)) ∙ 22sin( 393deg) = 06
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(378deg)) ∙ 17sin( 378deg) = 05
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3501 psi minus 50 psi) = 17480 psi
Eprime 15890 psi
Kprime = = = 8828 psi [3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Layer 3
Layer 4
Ground Stiffness and Soil Moduli
Layer 1
Values of qt and fs need to be in MPa
MR = (146(241 MPa)053 + 1355(012 MPa)14 + 236)244 = 3416 MPa = 49575 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1172 kPa m Vs (mfrasls) = [101 ∙ log(24139 kPa) minus 114]167 ∙ (100 ∙ ) = 2745
24139 kPa s
Vs = 9012 fts
29
γt 1204 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 37 ∙ (9012 ) = 304 ∙ 106psf = 21000 psi s
Layer 2
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3500 psi minus 100 psi) = 17480 psi
Dprime 17480 psi Eprime = = = 15890 psi
11 11
Eprime 15890 psi Kprime = = = 8828 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(241 MPa)053 + 1355(012 MPa)14 + 236)244 = 3415 MPa = 49562 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1172 kPa m Vs (mfrasls) = [101 ∙ log(24133 kPa) minus 114]167 ∙ (100 ∙ ) = 2745
24133 kPa s
Vs = 9012 fts
γt 1204 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
30
2 ft Gmax = ρt ∙ Vs
2 = 37 ∙ (9012 ) = 304 ∙ 106psf = 21000psi s
Layer 3
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (1500 psi minus 190 psi) = 7405 psi
Dprime 7405 psi Eprime = = = 6732 psi
11 11
Eprime 6732 psi Kprime = = = 3740 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(103 MPa)053 + 1355(008 MPa)14 + 236)244 = 1506 MPa = 21854 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 827 kPa m Vs (mfrasls) = [101 ∙ log(10343 kPa) minus 114]167 ∙ (100 ∙ ) = 2611
10343 kPa s
Vs = 8566 fts
γt 1172 pcf slug ρt = = = 36
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 36 ∙ (8566 ) = 27 ∙ 106psf = 19000 psi s
Layer 4
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (1200 psi minus 244 psi) = 5878 psi
31
Dprime 5878 psi Eprime = = = 5343 psi
11 11
Eprime 5343 psi Kprime = = = 2969 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(83 MPa)053 + 1355(005 MPa)14 + 236)244 = 165 MPa = 16901 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 483 kPa m Vs (mfrasls) = [101 ∙ log(8274 kPa) minus 114]167 ∙ (100 ∙ ) = 2243
8274 kPa s
Vs = 7360 fts
γt 1121 pcf slug ρt = = = 35
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 35 ∙ (7360 ) = 389 ∙ 106psf = 13000 psi s
32
2122 Example 2 Direct CPT Methods for Geoparameters on Clay
Several geoparameters need to be determined based on the given CPT data collected for the South
Abutment (Figure 15) The groundwater table (GWT) was measured at 60 feet Determine all the
geoparameters found in Table 5 at depths of 0 feet to 42 feet All sand layers can be assumed ldquodrainedrdquo
ν = 02 All clay layers can assume ν = 049
33
Figure 15 CPT data from Minnesota for example problem 2
34
Table 5 Geoparameters
Symbol Parameter Symbol Parameter
γt Soil total unit weight Eʹ Drained Youngrsquos modulus
Ic CPT material index Kʹ Bulk modulus
SBT Soil behavior type (SBT) MR Resilient modulus
ʹ σp Preconsolidation stress Gmax Small-strain shear modulus
YSR Yield stress ratio su Undrained shear strength
ϕʹ Effective friction angle cv Coefficient of consolidation
Ko Lateral stress coefficient k Hydraulic conductivity
Dʹ Constrained modulus
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
35
Figure 16 Soil layers using ldquorules of thumbrdquo
Unit weight of water γw = 6224 pcf
Layer 1
From Figure 16 fs = 13 psi taken as a representative value of the layer
13 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1179 pcf
145 psi
Layer 2
From Figure 16 fs = 12 psi taken as a representative value of the layer
12 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 3
From Figure 16 fs = 20 psi taken as a representative value of the layer
36
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
Layer 4
From Figure 16 fs = 2 psi taken as a representative value of the layer
2 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1004 pcf
145 psi
CPT Material Index
Layer 1
From Figure 10 fs = 13 psi and qc = 3000 psi
qt = qc + u2 ∙ (1 minus a) = 3000 psi + 0 psi ∙ (1 minus 08) = 3000 psi
σvo = γt ∙ 2 feet = 1179 pcf ∙ 2 feet = 235 psf = 16 psi
fs 13 psi Fr() = 100 ∙ = 100 ∙ = 043
(qt minus σvo) (3000 psi minus 16 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 16 psi minus 0 psi = 16 psi
(qt minus σvo)σatm (3000 psi minus 16 psi )145 psi = = Qtn prime )n (16 psi145 psi)n (σvoσatm
prime σvo 16 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(043)]2
37
Qtn = 4126 n = 032 lt 10 Ic = 121 therefore sand
Layer 2
From Figure 10 fs = 12 psi and qc = 250 psi
qt = qc + u2 ∙ (1 minus a) = 250 psi + 10 psi ∙ (1 minus 08) = 252 psi
σvo = 235 psf + γt ∙ 30 feet = 235 psf + 1172 pcf ∙ 30 feet = 3751 psf = 260 psi
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 53
(q minus σ ) (252 psi minus 260 psi) t vo
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 260 psi minus 0 psi = 260 psi
(qt minus σvo)σatm (250 psi minus 260 psi )145 psi = = Qtn prime (σvoσatm)n (260 psi145 psi)n
38
prime σvo 260 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(53)]2
Qtn = 87 n = 10 le 10 Ic = 32 therefore clay
Layer 3
From Figure 10 fs = 20 psi and qc = 4000 psi
qt = qc + u2 ∙ (1 minus a) = 4000 psi + 8 psi ∙ (1 minus 08) = 40016 psi
σvo = 3751 psf + γt ∙ 2 feet = 3751 psf + 1219 pcf ∙ 2 feet = 3995 psf = 277 psi
fs 16 psi Fr() = 100 ∙ = 100 ∙ = 054
(qt minus σvo) (30016 psi minus 277 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 277 psi minus 0 psi = 277 psi
(qt minus σvo)σatm (30016 minus 277)145 psi = = Qtn prime (σvoσatm)n (277 psi145 psi)n
39
primeσvo 277 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(054)]2
Qtn = 1962 n = 052 le 10 Ic = 15 therefore sand
Layer 4
From Figure 10 fs = 2 psi and qc = 250 psi
qt = qc + u2 ∙ (1 minus a) = 250 psi + 0 psi ∙ (1 minus 08) = 250 psi
= 3994 psf + γt ∙ 8 feet = 3994 psf + 1004 pcf ∙ 8 feet = 4798 psf = 333 psi σvo
fs 2 psi Fr() = 100 ∙ = 100 ∙ = 092
(qt minus σvo) (250 psi minus 333 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 333 psi minus 0 psi = 333 psi
(qt minus σvo)σatm (250 psi minus 333 psi )145 psi = = Qtn prime (σvoσatm)n (333 psi145 psi)n
40
prime σvo 333 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(092)]2
Qtn = 65 n = 10 lt 10 Ic = 29 therefore clayey silt
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic Qtn and Fr the first layer is defined as a ldquoDrained Sandrdquo from Figure 17
41
Figure 17 Soil layer 1 using SBT method
Layer 2
Based on values of Ic Qtn and Fr the second layer is defined as a ldquoUndrained Clayrdquo from Figure
18
42
Figure 18 Soil layer 2 using SBT method
43
Layer 3
Based on values of Ic Qtn and Fr the third layer is defined as a ldquoDrained Sandrdquo from Figure 19
Figure 19 Soil layer 3 using SBT method
44
Layer 4
Based on values of Ic Qtn and Fr the fourth layer is defined as a ldquoUndrained Silty Mixrdquo from
Figure 20
Figure 20 Soil layer 4 using SBT method
Effective Stress Friction Angle
Layer 1 (sand)
45
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(4126) = 464deg
Layer 2 (clay)
qt minus σvo
ϕprime(deg) = 295deg ∙ Bq0121 ∙ [0256 + 0336 ∙ Bq + log ( )]
primeσvo
Where
(u2 minus uo) (10 psi minus 0 psi) Bq = = = 004
(qt minus σvo) (252 psi minus 260 psi)
252 psi minus 260 psi ϕprime(deg) = 295deg ∙ 00440121 ∙ [0256 + 0336 ∙ 0044 + log ( )]
260 psi
ϕprime(deg) = 245deg
Layer 3 (sand)
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(1962) = 428deg
Layer 4 (clay)
qt minus σvo
ϕprime(deg) = 295deg ∙ Bq0121 ∙ [0256 + 0336 ∙ Bq + log ( )]
primeσvo
Where
(u2 minus uo) (5 psi minus 0 psi) Bq = = = 002
(qt minus σvo) (251 psi minus 333 psi)
252 psi minus 333 psi ϕprime(deg) = 295deg ∙ 0020121 ∙ [0256 + 0336 ∙ 002 + log ( )]
333 psi
ϕprime(deg) = 265deg
Stress History
46
Layer 1
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (12frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(3000 psi minus 33)072 = 1051 psi
prime σp 1051 psi YSR = = = 642
primeσvo 16 psi
Layer 2
028 028
prime m = 1 minus = 1 minus = 099 25 25 1 + (
Icfrasl ) 1 + (32frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(252 psi minus 260)099 = 734 psi
prime σp 734 psi YSR = = = 28
primeσvo 260 psi
Layer 3
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + ( frasl ) 1 + (15frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(4002 psi minus 277)072 = 1288 psi
47
prime σp 1288 psi YSR = = = 46
primeσvo 277 psi
Layer 4
028 028
prime m = 1 minus = 1 minus = 097 25 25 Ic 1 + ( frasl ) 1 + (29frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(250 psi minus 333)097 = 626 psi
prime σp 626 psi YSR = = = 19
primeσvo 333 psi
Lateral Stress Coefficient
Layer 1
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(464deg)) ∙ 642sin( 464deg) = 56
Layer 2
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(245deg)) ∙ 28sin( 245deg) = 090
Layer 3
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(428deg)) ∙ 46sin( 428deg) = 091
Layer 4
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(266deg)) ∙ 19sin( 266deg) = 073
48
Ground Stiffness and Soil Moduli
Layer 1
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3000 psi minus 16 psi) = 14991 psi
Dprime 14991 psi Eprime = = = 13629 psi
11 11
Eprime 13629 psi Kprime = = = 7571 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(207 MPa)053 + 1355(009 MPa)14 + 236)244 = 2816 MPa = 40873 psi
fs
03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 896 kPa m Vs (mfrasls) = [101 ∙ log(20685 kPa) minus 114]167 ∙ (100 ∙ ) = 2564
20685 kPa s
Vs = 8412 fts
γt 1179 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
2 ft 2 Gmax = ρt ∙ Vs = 37 ∙ (8412 ) = 260 ∙ 106psf = 17992 psi
s
Layer 2
49
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (252 psi minus 260 psi) = 11298 psi
Dprime 11298 psi Eprime = = = 10271 psi
11 11
Eprime 10271 psi Kprime = = = 17118 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(049))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(17 MPa)053 + 1355(008 MPa)14 + 236)244 = 443 MPa = 64309 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 827 kPa m Vs (mfrasls) = [101 ∙ log(17375 kPa) minus 114]167 ∙ (100 ∙ ) = 2646
17375 kPa s
Vs = 8680 fts
γt 1172 pcf slug ρt = = = 36
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 36 ∙ (8680 ) = 274 ∙ 106psf = 19036 psi s
Layer 3
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (4002 psi minus 277 psi) = 19869 psi
Dprime 19869 psi Eprime = = = 18063 psi
11 11
50
Eprime 18063 psi Kprime = = = 10035 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(276 MPa)053 + 1355(014 MPa)14 + 236)244 = 4017 MPa = 58309 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1379 kPa m Vs (mfrasls) = [101 ∙ log(27591 kPa) minus 1379]167 ∙ (100 ∙ ) = 2854
27591 kPa s
Vs = 9363 fts
γt 121 pcf slug ρt = = = 38
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 38 ∙ (9363 ) = 33 ∙ 106psf = 23050 psi s
Layer 4
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (250 psi minus 333 psi) = 1088 psi
Dprime 1088 psi Eprime = = = 989 psi
11 11
Eprime 989 psi Kprime = = = 16490 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(049))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
51
MR = (146(17 MPa)053 + 1355(001 MPa)14 + 236)244 = 360 MPa = 52189 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 138 kPa m Vs (mfrasls) = [101 ∙ log(17238 kPa) minus 114]167 ∙ (100 ∙ ) = 1545
17238 kPa s
Vs = 5069 fts
γt 1004 pcf slug ρt = = = 31
ga 322 ftfrasls2 ft3
2 ft 2 Gmax = ρt ∙ Vs = 31 ∙ (5069 ) = 80 ∙ 105psf = 5565 psi
s
Undrained Shear Strength
Layer 2
qt minus σv0 250 psi minus 26 psi su = = = 187 psi
12 12
Layer 4
qt minus σv0 250 psi minus 333 psi su = = = 181 psi
12 12
Coefficient of Consolidation
52
Example dissipation data are shown in Figure 21 Example calculations are provided
Figure 21 Dissipation t50 data
Layer 1
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (7197)075
cv = = = 359 cmsec 1 sec t50
Layer 2
53
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (94392)075
cv = = = 0008 cmsec 3000 sec t50
Layer 3
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (8303)075
cv = = = 665 cmsec 06 sec t50
Layer 4
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (26714)075
cv = = = 087 cmsec 11 sec t50
Hydraulic Conductivity
Layer 1
cm 1 psi 359 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 10 ∙ 10minus4 cmsec Dprime 14984 psi
Layer 2
cm 1 psi 0008 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 73 ∙ 10minus7 cmsec Dprime 11379 psi
Layer 3
cm 1 psi 665 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 20 ∙ 10minus5 cmsec Dprime 148650 psi
54
Layer 4
cm 1 psi 087 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 18 ∙ 10minus4 cmsec Dprime 10861 psi
55
CHAPTER 3 DIRECT CPT METHOD FOR SHALLOW
FOUNDATIONS
31 PROCEDURE
Shallow foundation analysis is typically done in a two-part traditional procedure The traditional
techniques are no longer required as a direct CPT method for square rectangular and circular shallow
footings is available (Figure 22) This process has the soil types grouped into four main categories sands
silts fissured clays and intact clays When determining soil types for each design it is believed footings
on sands and silts act in a fully drained manner while intact clays act in an undrained manner under
conditions of constant volume In order to determine the vertical stress-displacement-capacity of
square rectangular and circular shallow footings follow the steps provided towards the solution given
by Equation 27
Figure 22 Direct CPT method for shallow foundations
Equation 27 may be used to calculate all footing stresses from zero to the bearing capacity (qmax) To
calculate qmax for a sized footing of width (B) length (L) and thickness (t) follow steps 1 through 7
provided The settlement (s) can be determined after the calculation of qmax by simply rearranging
56
Equation 27 where the allowable stress (qallow) is defined as qmax divided by the factor of safety (FS) For
shallow footings a FS value of 3 is commonly used in geotechnical engineering
120782120787 minus120782120785120786120787 119956 119923 119954119950119938119961 = 119945119956 ∙ 119954119957119951119942119957 ∙ (119913
) ∙ (119913
) 27 119950119938119961
2 119902119898119886119909 0345 1 frasl119865119878 119871 119904 = 119861 ∙ [ ∙ ∙ ( ) ] 28
119861 ℎ119904 119902119905119899119890119905
Where hs = the foundation soil formation parameter qtnet = the net corrected cone tip resistance
311 Step 1 Estimating Footing Dimensions
In Minnesota frost heave can have devastating effect on a shallow foundations It is common practice to
place a foundation bearing elevation below the expected maximum frost depth (roughly 45 to 6 feet
below ground level) With this assumption estimate a footing size (B x L) for design to obtain
representative data from CPT roughly 15middotB below the foundation depth (Df) as shown in Figure 23 The
CPT data collected will consist of
cone tip resistance (qc)
measured porewater pressure acting behind the cone tip (u2) as shown in Figure 24
sleeve friction (fs)
57
Figure 23 Direct CPT method introduction
Figure 24 Differentiation of porewater pressure measurement locations (Lunne et al 1997)
312 Step 2 Soil Characterization
The soil behavior type (SBT) and CPT material index (Ic) govern the value of the formation factor hs The
first step is following the steps in Soil Unit Weight to determine soil layering and γt values for each
layer To determine a representative unit weight (γsoil) for all the layers in the range of Df to 15∙B below
58
Df the user will need to use their engineering judgement on the definition of ldquorepresentative unit
weightrdquo This single value of unit weight is used in further calculations such as the total vertical soil stress (σvo) from Equation 29 and effective vertical stress (σvo) from Equation 30 Values of σvo and σvo
will need to be calculated at 15middotB below Df
120648119959119952 = sum(120632119956119952119946119949 ∙ 119963) 29
prime 120648119959119952 = 120648119959119952 minus 119958119952 30
Where z = Df +15middotB uo=γwatermiddot(z-zw)
The qc will need to be corrected using Equation 31 in the case of fine-grained soils that develop excess porewater pressure during cone penetration These values will be used in the following calculations
119954119957 = 119954119940 + 119958120784 ∙ (120783 minus 119938) 31
Where 119860119899 a = cone area ratio = eg MnDOT commonly uses a = 08 119860119888
119860119899 = cross-sectional area of load cell or shaft 119860119888 = projected area of the cone
The cone area ratio is determined based on the type of piezocone tip used during in-situ field testing
Manufacturer specifications should provide the measured net area ratio (a) for the particular cone
penetrometer as determined by calibration in a pressurized triaxial chamber
313 Step 2a Foundation soil formation parameter
With the representative value γsoil continue to follow the steps in CPT Material Index through Soil
Behavior Type (SBT) to better define the type of soil at depth 15∙B below Df Use the value of Ic
calculated at depth 15∙B below Df to calculate hs with Equation 36 This parameter is based on the soil
type with typical values shown in Figure 25 The data on silts and sands are considered fully drained
whereas the fissured clay subset may be partially drained to undrained
59
23 ℎ119904 = 28 minus
119868119888 15 32
1+( ) 24
Figure 25 Foundation soil formation parameter hs versus CPT material index Ic (Mayne 2017)
314 Step 3 Soil elastic modulus and Poissonrsquos ratio
Details about the soil such as its elastic modulus (Es) and Poissonrsquos ratio (ν) will be needed for further
calculations A representative value for the Es can be determined from in-situ field tests Values of ν can
be assigned as 02 for drained sands and as 05 for undrained loading cases involving clays (Jardine et al
1985 Burland 1989)
315 Step 4 Net cone tip resistance
Calculate qtnet the mean value of net cone tip resistance 15middotB below the foundation bearing elevation
using Equation 33
33 119902119905119899119890119905 = 119902119905 minus 120590119907119900
60
316 Step 5 Bearing capacity of the soil
Use the assumed B and L calculated hs and qtnet to determine the soils bearing capacity (qmax) from
Equation 34 Use Table 6 to calculate qmax by using the maximum allowable settlement ratio (sB)max
correlating to the soil type If hs is in between the given values interpolate to acquire (sB)max
Table 6 Bearing capacity defined by soil type
Type of Soil hs (sB)max
Clean Sands 058 12
Silts 112 10
Fissured Clays 147 7
Intact Clays 270 4
05 minus0345 119904 119871 119902119898119886119909 = ℎ119904 ∙ 119902119905119899119890119905 ∙ (119861
) ∙ (119861
) 34 119898119886119909
317 Step 6 Settlement
Settlement can be calculated directly using the results from Equation 35 A FS equal to 3 is common in
foundation engineering
2 119902119898119886119909 0345 1 frasl119865119878 119871 119904 = 119861 ∙ [ ∙ ∙ ( ) ] 35
119861 ℎ119904 119902119905119899119890119905
318 Step 7 Final Check
Check that the applied stress (q) is less than qmax using Equation 36 Repeat the process again with a
new B andor new L if q gt qmax
05 119904 ) 119902 = ℎ119904 ∙ 119902119905119899119890119905 ∙ (119861
minus0345 119871 ∙ ( )
11986136
61
32 EXAMPLE PROBLEMS
321 Example 3 Direct CPT Method on Sands
A footing size needs to be determined based on the given CPT data collected for the South Abutment
(Figure 27) The footing stress (q) was determined to be 8000 psf Estimate a footing size (B x L) and
determine the bearing capacity of the foundation (qmax) using the direct CPT method provided Also
determine the expected settlement based on the calculated bearing capacity
Figure 26 Diagram of footing profiles for Example 3
The soil elastic modulus determined from seismic CPT (SCPT) are shown in Table 7
Table 7 SCPT Results
Depth (feet) Es (tsf)
Bottom of layer
3 557
6 433
9 557
12 695
17 590
22 501
27 571
32 505
62
Figure 27 CPT data from Northern Minnesota
Solution
Step 1 Assume the footing is placed below the frost depth of 6 feet
Estimate L L = 50 feet = 600 inches
63
Estimate B B = 12 feet = 144 inches
Estimate footing thickness t t = 2 feet
Df = 6 feet = 72 inches
Df + 15middotB = 24 feet = 288 inches
Step 2 Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 28 Schematic of foundation design associated with CPT data
Layer 1
From Figure 28 fs = 20 psi taken as a representative value of the layer
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
64
Layer 2
From Figure 28 fs = 20 psi taken as a representative value of the layer
20psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
Layer 3
From Figure 28 fs = 8 psi taken as a representative value of the layer
8 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1134 pcf
145 psi
Using the unit weights calculated between Df and 15B determine a representative unit weight of the soil to calculate the total and effective soil stresses Based on layer 3 being the weakest supporting layer γsoil = 113 pcf
σvo = γsoil ∙ (119863119891 + 15 ∙ B) = 113 lbsfraslft3 ∙ 24 feet = 2721 psf = 189 psi
prime σvo = σvo minus uo = 2721 psf minus 0 psf = 2721 psf = 189 psi
Calculate the cone tip resistance between Df and Df + 15B below the foundation depth
qt = qc + u2 ∙ (1 minus a) = 1250 psi + 0 psi ∙ (1 minus 08) = 1000 psi
Step 2a CPT Material Index
Using Figure 28 the sleeve friction at 15B below the foundation depth is about 16 psi
65
fs 16 psi Fr() = 100 ∙ = 100 ∙ = 13
(qt minus σvo) (1250 psi minus 189 psi)
Iterate to solve for Ic Steps are not shown for brevity
(qt minus σvo)σatm (1250 psi minus 189 psi )145 psi = = Qtn prime (σvoσatm)n (189 psi145 psi)n
prime σvo 189 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Qtn = 703 n = 072 lt 10 Ic = 210
66
Figure 29 Soil type for example problem 3
Use the Ic value to determine the foundation soil formation parameter
Calculate the foundation soil formation parameter The tip resistance is about 1250 psi at 24 feet and
the cone area ratio was determined to be 08 from information provided by the manufacturer
23 23 hs = 28 minus = 28 minus = 078 = SandSilt 15 15 Ic 210
1 + ( ) 1 + ( ) 24 24
67
Step 3 Determine representative values of soil elastic modulus from soil testing and Poissons ratio The
soil elastic modulus was taken as the average value between depths of Df and 15middotB (6 feet and 24 feet)
433 + 557 + 695 + 590 + 501 Es = = 555 tsf = 708 psi
5
ldquoDrained sandsiltrdquo gives a ν = 020
Step 4 Calculate the net cone tip resistance
qtnet = qt minus σvo = 1250 psi minus 189 psi = 12311 psi
Step 5 Calculate the bearing capacity of the sand
In drained sandssilts the bearing capacity is taken as the stress when (sB) = 011 (or 11 foundation width)
minus0345 minus0345 05 s L 600 qmax = hs ∙ qtnet ∙ ( ) ∙ ( ) = 078 ∙ (9795) ∙ (011)05 ∙ ( )
B B 144
= 193 psi = 27860 psf qmax
Assuming a factor of safety (FS) of 3
qmax = 645 psi = 9287 psf
3
Step 6 Calculate settlement
68
119902119898119886119909 0345 2
1 frasl119865119878 L 119904 = 119861 ∙ [ ∙ ∙ ( ) ]
ℎ119904 119902119905119899119890119905 B
2 0345 1 645 psi 600 inches s = 144 inches ∙ [ ∙ ∙ ( ) ] = 18 inches
078 12311 psi 144 inches
Step 7 Determine if q gt qmax
q = 8000 psf lt 9287psf = qmax3
69
CHAPTER 4 DIRECT CPT METHOD FOR DEEP FOUNDATIONS
41 INTRODUCTION
The axial compression capacity (Qtotal) for a single pile includes a side component (Qside) end bearing
component (Qbase) and pile weight (Wpile) as shown in Figure 30 Since piles commonly push through
several layers a summation of the unit side frictions acting on the pile segments must be considered
over the length of the pile While the examples shown in this Guide resemble hand calculations
computer software is more efficient Programming the procedure is possible and represents a practical
method of designing deep foundations However commercial software is frequently available and is
MnDOTrsquos most common method of designing deep foundations
There are upwards of 40 different direct CPT methods that have been developed over the past five
decades to determine a piles axial compression capacity Many of the earliest methods relied on hand-
recorded information where qc data from mechanical CPTs would be collected at 20 cm intervals
whereas the direct CPT method uses scaled penetrometer readings via specified algorithms to obtain
the pile unit side friction and unit end bearing The method that will be used for deep foundation design
is the Modified UniCone method which uses all three readings of the electronic piezocone (qt fs and u2)
while addressing a variety of pile foundation types
Figure 30 Direct CPT evaluation of axial pile capacity
70
The modified UniCone Method is based upon a total of 330 pile load tests (three times the original
Unicone database) that were associated with SCPTu data Originally the UniCone method provided
approximate soil classification in five groups via a chart of qE vs fs (Figure 31) where qE = qt -u2 Later
using the modified approach with a larger data set provided soil sub classifications as shown in Figure
32 This new 9-zone normalized soil behavior type is determined using CPT data in combination with the
CPT Material Index
Figure 31 UniCone Method soil behavior type using CPT (Mayne 2017)
Figure 32 Modified UniCone Method soil behavior type using CPT (Mayne 2017)
71
42 MODIFIED UNICONE METHOD
In order to determine the axial pile capacity using the modified method the first step requires the
determination of geoparameters as shown in Direct CPT Method for Soil Characterization Once the
soil unit weight and CPT Material index are determined for each soil layer then the effective cone
resistance pile unit side friction (fp) and pile end bearing resistance (qb) can be determined
421 Step 1
Work through the steps provided in CPT Method for Soil Characterization until each soil layer is
defined by its CPT material index and soil behavior type using Figure 6 After these steps have been
completed continue to Step 2 to determine qE qb and fp
422 Step 2
Once qt is determined for each layer the effective cone resistance can be calculated using Equation 37
119902119864 = 119902119905 minus 1199062 37
Where (a) qE is the specific value at each elevation along the pile sides for determining fp and (b) at the bottom of the pile qE is averaged in the vicinity of the pile tip from the tip bearing elevation to about one diameter beneath the tip for determining qb
Using CPT material index and qE the pile end bearing resistance is calculated using Equation 38
38 119902119887 = 119902119864 ∙ 10(0325∙119868119888minus1218)
The pile unit side friction is obtained from qE and Ic
39 119891119901 = 119902119864 ∙ 120579119875119879 ∙ 120579119879119862 ∙ 120579119877119860119879119864 ∙ 10(0732∙119868119888minus3605)
Where θPT = coefficient for pile type (084 for bored 102 for jacked 113 for driven piles) θTC = coefficient for loading direction (111 for compression and 085 for tension) θRATE = rate coefficient applied to soils in SBT zone 1 through 7 (109 for constant rate of penetration test and 097 for maintained load tests)
72
423 Step 3
Due to piles extending through multiple layers the unit side components acting on various pile
segments would need to be summed if not using the direct CPT method Since CPT calculates data at
regular intervals of 2 cm to 5 cm along the sides of the pile the average fp in each layer can be used
directly in Equation 40 to obtain the shaft capacity
119876119904119894119889119890 = 119891119901 ∙ 119860119904 40
Where fp = average pile side friction along pile length from eqn 39 As = πmiddotdmiddotH d = pile diameter and H = length embedded below grade
The base capacity for a pile in compression loading is given by Equation 41 For piles in tension (or uplift)
Qbase can be taken as 0
= 119902119887 ∙ 119860119887 41 119876119887119886119904119890 Where
qb = end bearing resistance from eqn 38 Ab = πmiddotd24 (area of a circular pile)
424 Step 4
The final step is to calculate the axial pile capacity using Equation 42
119876119905119900119905119886119897 = 119876119904119894119889119890 + 119876119887119886119904119890 minus 119882119901119894119897119890 42
43 AXIAL PILE DISPLACEMENTS
Movement of pile foundations can be assessed using elastic continuum theory which has been
developed using finite element analyses boundary elements and analytical closed-form solutions In
the case of piles passing through several soil layers the elastic solution can be used by stacking pile
segments (each with its own stiffness) as represented by soil Youngrsquos modulus The use of software is
recommended for pile groups Several available programs such as DEFPIG GROUP and PIGLET can
handle pile groups under axial and lateralmoment loading
73
44 EXAMPLE PROBLEMS
441 Example 5 Direct CPT Methods Axial Pile Capacity
Several piles need to be placed beneath the edge of a building Use the CPT data collected for this site
shown in Figure 34 to determine the axial capacity for one of the piles Assume round steel driven piles
will be used with a diameter of 1275 inches wall thickness of 025 inches and lengths of 80 feet
Concrete will be used to fill the piles To solve for all geoparameters use the Direct CPT Method for Soil
Characterization section All sand layers can be assumed ldquodrainedrdquo with ν = 02
Figure 33 Deep foundation end bearing pile diagram
74
Figure 34 CPT data from Minnesota for example problem 5
75
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 35 Soil layers using ldquorules of thumbrdquo for pile capacity example using direct CPT method
Layer 1
From Figure 35 fs = 18 psi taken as a representative value of the layer
76
18 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1210 pcf
145 psi
Layer 2
From Figure 35 fs = 12 psi taken as a representative value of the layer
12psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 3
From Figure 35 fs = 20 psi taken as a representative value of the layer
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
CPT Material Index
Layer 1
From Figure 35 fs = 18 psi and qc = 3000 psi
qt = qc + u2 ∙ (1 minus a) = 3000 psi + 20 psi ∙ (1 minus 08) = 3004 psi
∙ 4 feet = 1219 pcf ∙ 4 feet = 4877 psf = 34 psi σvo = γt
fs 18 psi Fr() = 100 ∙ = 100 ∙ = 060
(qt minus σvo) (3004 psi minus 34 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
77
uo = 0 psi
prime σvo = σvo minus uo = 34 psi minus 0 psi = 34 psi
(qt minus σvo)σatm (3004 psi minus 34 psi )145 psi = = Qtn prime )n (σvoσatm (34 psi145 psi)n
prime σvo 34 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(060)]2
Qtn = 3593 n = 038 lt 10 Ic = 14 (ie sand)
Layer 2
From Figure 35 fs = 12 psi and qc = 500 psi
qt = qc + u2 ∙ (1 minus a) = 500 psi + 40 psi ∙ (1 minus 08) = 508 psi
= 4838 psf + γt ∙ 45 feet = 4838 psf + 1172 pcf ∙ 45 feet = 5756 psf = 400 psi σvo
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 260
(qt minus σvo) (508 psi minus 400 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
78
uo = 0 psi
prime σvo = σvo minus uo = 400 psi minus 0 psi = 400 psi
(qt minus σvo)σatm (508 psi minus 400 psi )145 psi = = Qtn prime (σvoσatm)n (400 psi145 psi)n
prime σvo 400 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(260)]2
Qtn = 117 n = 10 le 10 Ic = 29 (ie clayey silt)
Layer 3
From Figure 35 fs = 20 psi and qc = 5000 psi
qt = qc + u2 ∙ (1 minus a) = 5000 psi + 0 psi ∙ (1 minus 08) = 5000 psi
= 5756 psf + γt ∙ 6 feet = 5756 psf + 1219 pcf ∙ 6 feet = 6488 psf = 451 psi σvo
fs 20 psi Fr() = 100 ∙ = 100 ∙ = 040
(qt minus σvo) (5000 psi minus 451 psi)
79
Step 2 Iterate to solve for Ic Steps are not shown for brevity
prime σvo = σvo minus uo = 451 psi minus 0 psi = 451 psi
(qt minus σvo)σatm (5000 psi minus 451 psi )145 psi = = Qtn prime )n (σvoσatm (451 psi145 psi)n
prime σvo 451 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(040)]2
= 1801 n = 06 lt 10 = 15 (ie sand) Qtn Ic
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic (14) Qtn (359) and Fr (06) the first layer is defined as a ldquoGravelly Sandrdquo
from Figure 36
80
Figure 36 Soil layer 1 using SBT method
81
Layer 2
Based on values of Ic (29) Qtn (117) and Fr (26) the second layer is defined as a ldquoSilty Mixrdquo
from Figure 37
Figure 37 Soil layer 2 using SBT method
82
Layer 3
Based on values of Ic (15) Qtn (180) and Fr (04) the third layer is defined as a ldquoSandrdquo from
Figure 38
Figure 38 Soil layer 3 using SBT method
Effective Cone Resistance
Layer 1
83
qE = qt minus u2 = 3004 psi minus 20 psi = 2984 psi
Layer 2
qE = qt minus u2 = 508 psi minus 40 psi = 468 psi
Layer 3
qE = qt minus u2 = 5000 psi minus 0 psi = 5000 psi
Pile End Bearing Resistance
Layer 3
= qE ∙ 10(0325∙Icminus1218) qb = 5000 psi ∙ 10(0325∙15minus1218) = 9085 psi
Pile Unit Side Friction
Layer 1
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 2984 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙14minus3605) = 99 psi
Layer 2
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 468 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙29minus3605) = 211 psi
Layer 3
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 5000 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙15minus3605) = 202 psi
84
Axial Pile Capacity
The side capacity is based on pile design as shown in Figure 39
Layer 1
Qside = fp ∙ As = 99 psi ∙ (π ∙ d ∙ L) = 99 psi ∙ (π ∙ 1275 in ∙ 48 in) = 19064 lb
Layer 2
Qside = fp ∙ As = 211 psi ∙ (π ∙ d ∙ L) = 211 psi ∙ (π ∙ 1275 in ∙ 588 in) = 457122 lb
Layer 3
Qside = fp ∙ As = 202 psi ∙ (π ∙ d ∙ L) = 202 psi ∙ (π ∙ 1275 in ∙ 240 in) = 58170 lb
Figure 39 Soil layering compared to pilings
85
End Bearing Resistance in Layer 3
d2 1275 in2
Qbase = qb ∙ Ab = 9085 psi ∙ (π ∙ ) = 9085 psi ∙ (π ∙ ) = 115932 lb 4 4
Wp = (γsteel ∙ Asteel ∙ Depth) + (γsteel ∙ Aconc ∙ Depth)
Wp = (490 pcf ∙ 003 ft2 ∙ 55 ft) + (150 pcf ∙ 085 ft2 ∙ 55 ft) = 7724 lbs
Layer 3
Qtotal = ΣQside + Qblayer3 minus Wp
Qtotal = (19064 + 45713 + 58170) lbs + 115932 lbs minus 7724 lbs = 642563 lbs
= 642 kips Qtotal
Table 8 Summary of selected and calculated parameters
Layer L (in) qc
(psi)
fs
(psi)
Fr Ic Qtn qE
(psi)
qb
(psi)
Qbase
(lb)
fp
(psi)
Qside
(lb)
Qtotal
(kip)
1 48 3000 18 06 14 356 2984 NA NA 99 19064
2 588 500 12 26 29 117 468 NA NA 211 457122
3 240 5000 20 04 15 180 5000 9085 115932 202 58170
Total 115932 534355 642
86
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American Petroleum Institute (API) (1987) Planning designing and constructing fixed offshore platforms Recommended practice API-RP2A 18th edition Washington DC API
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Berardi R Jamiolkowski M and Lancellotta R (1991) Settlement of shallow foundations in sands selection of stiffness on the basis of penetration resistance Geotechnical Engineering Congress Vol 1 185-200 (Proc GSP-27 Boulder) Reston Virginia ASCE
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Briaud JL and Gibbens RM (1999) Behavior of five large spread footings in sand Journal of Geotechnical amp Geoenvironmental Engineering 125 (9) 787-797
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Brown DA Turner JP and Castelli RJ (2010) Drilled Shafts Construction Procedures and LRFD Design Methods Geotechnical Engineering Circular GEC 10 Report No FHWA NHI-10-016 Washington DC National Highway Institute US DOT
87
Burland J B (1973) Shaft Friction of Piles in Clay -A Simple Fundamental Approach Ground Eng 6(3) 30-42
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Canadian Geotechnical Society (1992) Canadian Foundation Engineering Manual Third Edition Vancouver British Columbia BiTech Publishers
Cerato AB and Lutenegger AJ (2007) Scale effects of shallow foundation bearing capacity on granular material Journal of Geotechnical amp Geoenvironmental Engineering 133 (10) 1192-1201
Chen BS-Y and Mayne PW (1996) Statistical relationships between piezocone measurements and stress history of clays Canadian Geotechnical Journal 33 (3) 488-498
Cho W and Finno RJ (2010_ Stress-strain responses of block samples of compressible Chicago glacial clays Journal of Geotechnical amp Geoenvironmental Engineering 136 (1) 178-188
Chung SF Randolph MF and Schneider JA (2006) Effect of penetration rate on penetrometer resistance in clay Journal of Geotechnical amp Geoenvironmental Engineering 132 (9) 1188-1196
Clausen CJF Aas PM and Karlsrud K (2005) Bearing capacity of driven piles in sand the NGI approach Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis Group
Consoli NC Schnaid F and Milititsky J (1998) Interpretation of plate load tests on residual soil site Journal of Geotecnical amp Geoenvironmental Engingeering 124 (9) 857-867
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DeBeer EE (1965) Bearing capacity and settlement of shallow foundations on sand Proc Symposium on Bearing Capacity and Settlement of Foundations 15-33 Duke University Durham NC
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Demers D and Leroueil S (2002) Evaluation of preconsolidation pressure and the overconsolidation ratio from piezocone tests of clay deposits in Quebec Canadian Geotechnical Journal 39 (1) 174-192
Eslaamizaad S and Robertson PK (1996) Cone penetration test to evaluate bearing capacity of foundations in sand 429-438 Proc CGCFG-49 Vol 1 St Johns Newfoundland
Eslami A Alimirzaei M Aflaki E and Molaabasi H (2017) Deltaic soil behavior classification using CPTu records Marine Georesources amp Geotechnology 35 (1) 62-79
88
Eslami A and Gholami M (2005) Bearing capacity analysis of shallow foundations from CPT data 1463-1466 Proc ICSMGE- 16 Vol 3 (Osaka) Millpress Rotterdam
Eslami A and Gholami M (2006) Analytical model for the ultimate bearing capacity of foundations from cone resistance Scientia Iranica 13 (3) 223-233
Eslami A and Fellenius BH (1997) Pile capacity by direct CPT and CPTu methods applied to 102 case histories Canadian Geotechnical Journal 34 (6) 886-904
Fellenius BH (2009) Basics of Foundation Design Vancouver BiTech Publishers
Fellenius BH and Altaee A (1994) Stress and settlement of footings in sand Vertical and Horizontal Deformations of Foundations and Embankments 2 1760-1773
Fellenius BH and Eslami A (2000) Soil profile interpreted from CPTu data Proc Geotechnical Engineering Conference Year 2000 Geotechnics Asian Inst of Technology Bangkok Thailand Available from wwwfelleniusnet
Finno R J and Chung C K (1992) Stress-strain-strength responses of compressible Chicago glacial clays Journal of Geotechnical Engineering 118 (10) 1607ndash1625
Fleming K Weltman A Randolph M and Elson K (2009) Piling Engineering Third Edition London Taylor amp Francis Group
Frank R and Magnan J-P (1995) Cone penetration testing in France national report Proceedings International Symposium on Cone Penetration Testing 3 (CPT95 Linkoumlping) Swedish Geotechnical Society Report 3(95) 147-156
Gavin K Adekunte A and OKelly B (2009) A field investigation of vertical footing response on sand Geotechnical Engineering 162 257-267
Hegazy YA (1998) Delineating geostratigraphy by cluster analysis of piezocone data PhD dissertation School of Civil amp Environmental Engineering Georgia Institute of Technology Atlanta GA
Holtz RD Kovacs WD and Sheehan TC (2011) An Introduction to Geotechnical Engineering 2nd Edition London Pearson Publishing
Hunt CE Pestana JM Bray JD and Riemer M (2002) Effect of pile driving on static and dynamic properties of soft clay Journal of Geotechnical amp Geoenvironmental Engineering 128 (1) 13-24
Jamiolkowski M Ladd CC Germaine JT and Lancellotta R (1985) New developments in field and lab testing of soils Proc ICSMFE-11 1 57-154
Jardine R Chow F Overy R and Standing J (2005) ICP design methods for driven piles in sands and clays London Thomas Telford Publishing
Jardine RJ Lehane BM Smith P and Gildea P (1995) Vertical loading experiments on rigid pad foundations at Bothkennar Geotechnique 45 (4) 573-597
89
Karlsrud K (2012) Prediction of load-displacement behavior and capacity of axially-loaded piles in clay based on analyses and interpretation of pile load test results PhD Dissertation Norwegian Univ Science amp Technology Trondheim Norway
Karlsrud K Clausen CJF and Aas PM (2005) Bearing capacity of driven piles in clay the NGI approach Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis
Karlsrud K Lunne T Kort DA and Strandvik S (2001) CPTU correlations for clays Proc 16th ICSMGE 2 683-702
Kimura T Kusakabe O and Saitoh K (1985) Geotechnical model tests of bearing capacity problems in a centrifuge Geotechnique 35 (1) 33-45
Knappett JA and Craig RF (2012) Craigs Soil Mechanics 8th Edition London Spon Press Taylor amp Francis
Kolk HJ and Velde EVD (1996) A reliable method to determine friction capacity of piles driven into clays Paper No 7993 Proc Offshore Technology Conference Houston
Kolk HJ Baaijens AE and Senders M (2005) Design criteria for pipe piles in silica sands Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis Group
Kulhawy FH (2004) On the axial behavior of drilled shafts Geosupport 2004 1-18
Kulhawy FH and Jackson CS (1989) Foundation Engineering Current Principles amp Practices 2 (GSP 22) Reston VA ASCE
Kulhawy FH and Mayne PW (1990) Manual for Estimating Soil Properties for Foundation Design EPRI Report EL-6800 Electric Power Research Institute Palo Alto 306 page Download from wwwepricom
Kulhawy FH Trautmann CH Beech JF ORourke TD and McGuire W (1983) Transmission Line Structure Foundations for Uplift-Compression Loading Report EL-2870 Palo Alto CA Electric Power Research Institute wwwepricom
Ku T and Mayne PW (2015) In-situ lateral stress coefficient (K0) from shear wave velocity measurements in soils Journal of Geotechnical amp Geoenvironmental Engineering 141 (12) 101061(ASCE) GT1943-56060001354
Ladd CC and DeGroot DJ (2003) Recommended practice for soft ground site characterization Soil amp Rock America 2003 3-57
Larsson R (1997) Investigations and Load Tests in Silty Soils SGI Report R-54 Linkoumlping Swedish Geotechnical Institute
Larsson R (2001) Investigations and Load Tests in Clay Till SGI Report R-59 Linkoumlping Swedish Geotechnical Institute
Lee J and Salgado R (2005) Estimation of bearing capacity of circular footings on sands based on cone penetration tests Journal of Geotechnical amp Geoenvironmental Engineering 131 (4) 442-452
90
Lehane BM (2003) Vertically loaded shallow foundation on soft clayey silt Geotechnical Engineering 156 (1) Thomas Telford London 17-26
Lehane BM (2013) Relating foundation capacity in sands to CPT qc Geotechnical amp Geophysical Site Characterization 1 London Taylor amp Francis Group
Lehane BM Doherty JP and Schneider JA (2008) Settlement prediction for footings on sand Deformational Characteristics of Geomaterials (IS-Atlanta) 1 Rotterdam Millpress
Lehane BM Li Y and Williams R (2013) Shaft capacity of displacement piles in clay using the CPT Journal of Geotechnical amp Geoenvironmental Engineering 139 (2) 253-266
Lehane BM Schneider JA and Xu X (2005) The UWA-05 method for prediction of axial capacity of driven piles in sand Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis
Low HE Lunne T Andersen KH Sjursen MA Li X and Randolph MF 2010 Estimation of intact and remoulded undrained shear strengths from penetration tests in soft clays Geotechnique 60 (11) 843-859
Lunne T Robertson PK and Powell JJM (1997) Cone Penetration Testing in Geotechnical Practice London RoutledgeBlackie Academic
Lutenegger AJ and DeGroot DJ (1995) Settlement of Shallow Foundations on Granular Soils Report Contract 6332 Task Order 4 submitted to Massachusetts Highway Dept by University of Mass-Amherst Amherst MA
Lutenegger AJ and Adams MT (2003) Characteristic load-settlement behavior of shallow foundations Proc FONDSUP 2 381-392
Mase T and Hashiguchi K (2009) Numerical analysis of footing settlement problem by subloading surface model Soils amp Foundations 49 (2) 207-220
Mayne PW (1987) Determining preconsolidation stress and penetration pore pressures from DMT contact pressures Geotechnical Testing Journal 10(3) 146-150
Mayne PW (1991) Determination of OCR in Clays by Piezocone Tests Using Cavity Expansion and Critical State Concepts Soils and Foundations 31 (2) 65-76
Mayne PW (2007) NCHRP Synthesis 368 on Cone Penetration Test Washington DC Transportation Research Board National Academies Press wwwtrborg
Mayne PW (2009) Geoengineering Design Using the Cone Penetration Test Richmond BC Canada ConeTec Inc
Mayne PW (2014) Keynote Interpretation of geotechnical parameters from seismic piezocone tests Proceedings 3rd Intl Symp on Cone Penetration Testing 47-73
Mayne PW (2016) Evaluating effective stress parameters and undrained shear strengths of soft-firm clays from CPT and DMT Australian Geomechanics Journal 51 (4) 27-55
91
Mayne PW (2017) Stress history of soils from cone penetration tests 34th Manual Rocha Lecture Soils amp Rocks Vol 40 (3) Saouml Paulo ABMS
Mayne PW Christopher B Berg R and DeJong J (2002) Subsurface Investigations -Geotechnical Site Characterization Publication No FHWA-NHI-01-031 Washington DC National Highway Institute Federal Highway Administration
Mayne PW Coop MR Springman S Huang A-B and Zornberg J (2009) Geomaterial Behavior and Testing Proc 17th Intl Conf Soil Mechanics amp Geotechnical Engineering 4 Rotterdam MillpressIOS Press
Mayne PW and Illingworth F (2010) Direct CPT method for footings on sand using a database approach Proc ISCPT-2 3 315-322
Mayne PW and Niazi F (2017) Recent developments and applications in field investigations for deep foundations Proceedings 3rd Bolivian Conference on Deep Foundations 1 141-160
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Mayne PW and Poulos HG (1999) Approximate displacement influence factors for elastic shallow foundation systems ASCE Journal of Geotechnical amp Geoenvironmental Engineering 125 (6) 453-460
Mayne PW Uzielli M and Illingworth F (2012) Shallow footing response on sands using a direct method based on cone penetration tests Full Scale Testing and Foundation Design (GSP 227) Reston Virginia ASCE
Mayne PW and Woeller DJ (2014) Direct CPT method for footings on sands silts and clays Geocharacterization for Modeling and Sustainability (GSP 234) 1983-1997
McClelland B (1974) Design of Deep Penetration Piles for Ocean Structures Journal Geot Eng Div 100(GT7) 705-747
Mesri G (1975) Discussion on new design procedure for stability of soft clays ASCE Journal of the Geotechnical Engineering Division 101(GT4) pp 409-412
Meyerhof GG (1956) Penetration tests and bearing capacity of cohesionless soils Journal of Soil Mechanics amp Foundations Division 82 (SM1) 1-19
Meyerhof GG (1974) General Report Penetration testing outside Europe Proceedings ESOPT-1 Vol 21 Swedish Geotechnical Society Stockholm
Meyerhof GG (1976) Bearing capacity and settlement of pile foundations Journal of Geotechnical Engineering Division 102(3) 197-228
Missiaen T Verhegge J Heirman K and Crobe P (2015) Potential of cone penetrating testing for mapping deeply buried palaeolandscapes in the context of archaeological surveys in polder areas Journal of Archaeological Science 55 174-187
92
Mlynarek Z Wierzbicki J Goglik S and Bogucki M (2014) Shear strength and deformation parameters of peat and gyttja from CPTu DMT and VST Proceedings 5th International Workshop on CPTu and DMT in soft clays and organic soils 193-210 Poznan Poland Polish Committee on Geotechnics Menard Polska Tensar Internationa
Nejaim PF Jannuzzi GMF and Danziger FAB (2016) Soil behavior type of the Sarapuiacute II test site Geotechnical amp Geophysical Site Characterization 5 1009-1014
Niazi FS and Mayne PW (2013) Cone penetration test based direct methods for evaluating static axial capacity of single piles Geotechnical and Geological Engineering 31 (4) 979-1009
Niazi FS and Mayne PW (2015) Enhanced Unicone expressions for axial pile capacity evaluation from piezocone tests Proc International Foundations Conference amp Equipment Expo RestonVA ASCE
Niazi FS and Mayne PW (2016) CPTu-based enhanced UniCone method for pile capacity Engineering Geology 212 21-34
Nowacki F Karlsrud K and Sparrevik P (1996) Comparison of recent tests on OC clay and implications for design Large Scale Pile Tests in Clay London Thomas Telford
ONeill MW (2001) Side resistance in piles and drilled shafts Journal of Geotechnical amp Geoenvironmental Engineering 127 (1) 3-16
Ouyang Z and Mayne PW (2018) Effective friction angle of clays and silts from piezocone Canadian Geotechnical Journal in press doiorg101139cgj-2017-0451
Paikowsky SG Canniff MC Lesny K Kisse A Amatya S and Muganga R (2010) LRFD Design and Construction of Shallow Foundations for Highway Bridge Structures NCHRP Report 651 Washington DC National Cooperative Highway Research Program Transportation Research Board
Pestana JM Hunt CE and Bray JD (2002) Soil deformation and excess pore pressure filed around a closed-ended pile Journal of Geotechnical amp Geoenvironmental Engineering 128 (1) 1-12
Poulos HG and Davis EH (1974) Elastic Solutions for Soil and Rock Mechanics New York Wiley amp Sons
Poulos HG and Davis E (1980) Pile Foundation Analysis amp Design New York John Wiley amp Sons
Powell JJM and Lunne T (2005) Use of CPTU data in clays and fine-grained soils Studia Geotechnica et Mechanica XXVII(3) 29-66
Randolph MF (2003) Rankine Lecture Science and empiricism in pile foundation design Geotechnique 53 (10) 847-875
Robertson PK (1990 1991) Soil classification using the cone penetration test Canadian Geotechnical Journal 27 (1) 151-158 Closure 28 (1) 176-178
Robertson PK (1991) Estimation of foundation settlements in sand from CPT Geotechnical Engineering Congress 2 Reston Virginia ASCE
93
Robertson PK (2009) Cone penetration testing a unified approach Canadian Geotechnical J 46 (11) 1337-1355
Robertson PK (2009a) Performance based earthquake design using the CPT Keynote Lecture International Conference on Performance-Based Design in Earthquake Geotechnical Engineering IS Tokyo Tsukuba Japan
Robertson PK (2009b) Interpretation of cone penetration tests a unified approach Canadian Geotechnical Journal 46 (11) 1337-1355
Robertson PK and Cabal KL (2007) Guide to cone penetration testing Second Edition Signal Hill California Gregg In-Situ
Robertson PK and Cabal KL (2015) Guide to Cone Penetration Testing for Geotechnical Engineering 6th Edition Signal Hill California Gregg Drilling amp Testing Inc
Schmertmann JH (1970) Static cone to compute static settlement over sand Journal of the Soil Mechanics and Foundations Division 95 (SM3) 1011-1043
Schmertmann JH (1978) Guidelines for cone penetration test performance and design Report FHWA-TS-78-209 Washington DC Federal Highway Administration
Schnaid F Wood WR Smith AKC and Jubb P (1993) Investigation of bearing capacity and settlements of soft clay deposits at Shellhaven Predictive Soil Mechanics London Thomas Telford
Schneider JA Xu X and Lehane BM (2008) Database assessment of CPT-based design methods for axial capacity of driven piles in siliceous sands J Geotechnical and Geoenvironmental Engineering 134 (9) 1227-1244
Semple D (1980) Recent Developments in the Design amp Construction of Piles London Institution of Civil Engineers
Senneset K Sandven R amp Janbu N (1989) Evaluation of soil parameters from piezocone tests In-Situ Testing of Soil Properties for Transportation Transportation Research Record 1235 24-37
Shahri AA Melehmir A and Juhlin C (2015) Soil classification analysis based on piezocone penetration test data Engineering Geology 189 32-47
Stuedlein AW and Holtz RD (2010) Undrained displacement behavior of spread footings in clay The Art of Foundation Engineering Practice GSP 198 Reston Virginia ASCE
Takesue K Sasao H and Matsumoto T (1998) Correlation between ultimate pile skin friction and CPT data Geotechnical Site Characterization 2 1177ndash1182
Tand KE Funegard EG and Briaud J-L (1986) Bearing capacity of footings on clay CPT method Use of In-Situ Tests in Geotechnical Engineering GSP 6 1017-1033
Tand KE Warden PE and Funegaringrd EG (1995) Predicted-measured bearing capacity of shallow footings on sand PISCPT 2 589-594
Thomas D (1968) Deep soundings test results and the settlement of spread footings on normally consolidated sands Geotechnique 18 (2) 472-488
94
Tomlinson MJ (1957) The adhesion of piles driven into clay soils Proc 4th Intl Conf on Soil Mechanics and Foundation Engineering 2 66-71
Tomlinson MJ (1986) Foundation Design amp Construction London Longman Scientific
Tomlinson MJ (1987) Pile Design and Construction Practice London Viewpoint Publication
Trofimenkov JG (1974) Penetration testing in the USSR Proceedings ESOPT-1 1 147-154
Tumay MT Hatipkarasulu Y Marx ER and Cotton B (2013) CPTPCPT-based organic material profiling Proc 18th Intl Conf on Soil Mechanics amp Geotechnical Engineering Paris 633-636
Uzielli M and Mayne PW (2011) Statistical characterization and stochastic simulation of load-displacement behavior of shallow footings Proc Georisk 2011 Geotechnical Risk Management amp Assessment (GSP 224) 672-679
Uzielli M and Mayne PW (2012) Load-displacement uncertainty of vertically-loaded shallow footings on sands and effects on probabilistic settlement estimation Georisk Assessment and Management of Risk for Engineered Systems and GeoHazards 6 (1) 50-69
Valsson SM (2016) Detecting quick clay with CPTu Proceedings 17th Nordic Geotechnical Meeting Reykajavik Iceland Challenges in Nordic Geotechnics
Van Dijk BFJ and Kolk HJ (2011) CPT-based design method for axial capacity of offshore piles in clay Frontiers in Offshore Geotechnics II 555-560
Vesić AS (1975) Bearing capacity of shallow foundations Foundation Engineering Handbook New York Van Nostrand Reinhold Publishers
Vesic AS (1977) NCHRP Synthesis of Highway Practice 42 Design of Pile Foundations Washington DC Transportation Research Board
Yu F and Yang J (2012) Base capacity of open-ended steel pipe piles in sand J Geotechnical and Geoenvironmental Engineering 138 (9) 1116-1128
Zawrzykraj P Rydelek P and Bakowska A (2017) Geoengineering properties of Eemian peats from Radsymin (central Poland) in the light of static cone penetration and dilatometer tests Engineering Geology 226 290-300
95
APPENDIX A
List of Figures
Figure A1 Conventional methods versus direct CPT approach to shallow foundation response A-3
Figure A2 Direct bearing capacity relationship for embedded square and strip footings situated on clean
sands (after Schmertmann 1978) A-6
Figure A3 Direct CPT bearing factor Rk as function of footing width B and embedment depth from finite
element analyses (Tand et al 1995) A-6
Figure A4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio and sand
consistency (Eslaamizaad amp Robertson 1996) A-7
Figure A5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp Salgado (2005) in
terms of footing size relative density and base settlement A-8
Figure A6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and normalized
cone tip resistance (after Eslami amp Gholami 2005) A-8
Figure A7 Direct relationship between ultimate bearing stress in clays and measured cone tip resistance
(Schmertmann 1978) A-10
Figure A8 Direct relationship between ultimate foundation bearing stress in clays and net cone tip
resistance (after Tand et al 1986) A-11
Figure A9 Measured load-displacements for each of the five large footings at TAMU (Briaud amp Gibbens
1994) sand site A-12
Figure A10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU (Briaud amp
Gibbens 1994) footings A-13
Figure A11 Characteristic stress versus square root normalized displacements for TAMU footings A-15
Figure A12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sand A-16
Figure A13 Summary of sand formation factors with corresponding CPT resistances A-19
Figure A14 Normalized foundation stresses vs square root of normalized displacement for spread
footings on sand (modified after Mayne et al 2012) A-20
Figure A15 Definitions of capacity from three types of observed foundation load test responses
(modified after Kulhawy 2004) A-21
Figure A16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footings A-22
A-1
Figure A17 Measured load vs displacement curves for 3 shallow foundations on clay at Baytown Texas
(Stuedlein amp Holtz 2010) A-25
Figure A18 Characteristic stress vs square root of normalized displacement for all three foundations at
Baytown site (Stuedlein amp Holtz 2010) A-26
Figure A19 Normalized foundation stress vs square root of normalized displacements A-27
Figure A20 Applied foundation stress vs CPT-calculated stress for 67 footings A-28
Figure A21 Applied footing stress vs CPT calculated stress on logarithmic scales A-29
Figure A22 Summary graph and equations of direct CPT method for evaluating the vertical A-30
Figure A23 Foundation soil formation parameter hs versus CPT material index Ic A-32
Figure A24 Bearing capacity ratio qmaxqnet versus CPT material index Ic A-33
Figure A25 Influence factor for rectangular foundations from elastic theory solution (Mayne amp
Dasenbrock 2017) A-34
List of Tables
Table A1 Direct CPT methods for bearing capacity of footings on clean sands A-4
Table A2 Direct CPT methods for bearing capacity of footings on clays A-9
Table A3 Summary of large footings soil conditions and reference sources of database A-17
Table A4 Sources for shallow foundation performace A-34
A-2
A1 INTERPRETATION OF SOIL PARAMETERS FROM CONE
PENETRATION TESTS
A11 Introduction
Geotechnical site characterization is an important first step towards the evaluation of
subsurface conditions and determination of soil layering geomaterial classification and the
evaluation of soil engineering parameters for the analysis and design of foundations retaining
walls tunnels excavations embankments and slope stability Increasing use of the electronic
cone penetrometer in highway applications is prominent in the USA because the results are
obtained much faster and less expensive than traditional methods that rely on rotary drilling
augering sampling and laboratory testing Moreover the cone penetration test (CPT) provides
at least three independent and continuous readings on soil behavior that are digitally recorded
and fully available at the end of the sounding As such the reliable and consistent
interpretation of CPT data is important since civil engineering works will best realize the
efficiency and economy of this technology in constructed facilities for county state and
interstate projects
A111 Cone Penetration Testing
The cone penetrometer is an electronically-instrumented steel probe that is vertically pushed into the
ground at constant rate of 20 mms (08 insec) using a hydraulic system The penetrometer is outfitted
with load cells pressure transducers and inclinometers that measure at least three readings
continuously with depth (a) cone tip resistance qt (b) sleeve friction fs and (c) porewater pressure u2
With the latter reading the device is called a piezocone thus the designation CPTu is used to indicate
porewater pressure Additional sensors are available that can be used to measure inclination
resistivity pH temperature shear wave velocity and other readings if desired
Figure A1 shows a schematic rendering of the cone penetration test that is performed in the
field per ASTM D 5778 procedures The hydraulic system is nominally rated at 180 kN (20 tons)
capacity and often mounted on a truck but can also be positioned on tracked vehicles or
anchored frames Figure A2 shows a MnDOT cone truck that uses the full dead weight of the
rig for the hydraulic reaction forces which are applied at the center of the vehicle The cab is
enclosed so that soundings may proceed during inclement weather and the data acquisition
system and operator are protected
Figure A 1 Setup and procedures for co ne penetration testing (CPT)
A-4
Figure A2 CPT truck-mounted rig used by MnDOT
A representative sounding taken at the I-35 test site just northeast of Saint Paul is presented in Figure
A3 Here the separate profiles of qt fs and u2 with depth are shown in side-by-side plots In the last
graph the hydrostatic porewater pressure (uo) due to the groundwater table is indicated by the blue
dashed line
These readings are used to interpret the soil layers types and properties of the ground as
discussed in the following sections
A-5
Figure A3 Representative CPT sounding at I-35 test site northeast of St Paul MN showing profiles of
(a) cone resistance (b) sleeve friction and (c) penetration porewater pressures
A112 Soil Parameter Interpretation
A variety of soil engineering parameters can be interpreted from CPT results on the basis of
theoretical frameworks analytical models or numerical simulations otherwise by empirical
methods based on correlations and statistics of databases Figure A4 shows a conceptual
utilization of the readings from the CPT for interpretation of geostratigraphy compressibility
flow characteristics and stress-strain-strength behavior of soils
Table A1 lists a selection of geoparameters that have been addressed for these purposes The
most common or useful relationships will be discussed in subsequent sections and reference is
made to other sources (Lunne et al 1997 Mayne 2007 Robertson amp Cabal 2015) for additional
details
A-6
Figure A4 Conceptual post-processing of CPT results for geoparameter evaluation
Table A1 List of common geoparameters interpreted from CPT data
Symbol Parameter Remarks Notes
SBT Soil behavior type (SBT)
Ic CPT material index
γt Unit weight
ρt Mass Density
σvo Overburden stress
u0 Hydrostatic pressure
A-7
σvo Effective stress
Table A1 Continued
Symbol Parameter Remarks Notes
σp Preconsolidation stress σp = 033 (qnet)m where m = fctn (Ic)
(or Effective Yield Stress) where stresses are in kPa
YSR Yield stress ratio YSR = σpσvo (formerly OCR)
c Effective cohesion intercept
ϕ Effective friction angle
su Undrained shear strength
D Constrained modulus
G0 Small-strain modulus
G Shear modulus
E Youngs modulus
cvh Coefficient of consolidation
K Bulk modulus
A-8
LI
K0 Lateral stress coefficient
k Hydraulic conductivity
Liquefaction index
A12 Geostratigraphy and Soil Behavioral Type (SBT)
In routine CPT soil samples are not normally taken and thus the measured stress friction andor
porewater pressure readings are used to infer the types of soils This can be accomplished using
three basic approaches
a Rules of Thumb or approximate guidelines for a quick visual assessment
b Soil Behavioral Type (SBT) Charts that are based on either the three readings (qt fs u2) net readings
including net cone resistance (qnet = (qt-σvo) effective cone resistance (qE = qt - u2) friction ratio (Rf =
100middotfsqt) and excess porewater pressures or using normalized CPT readings such as Q F Bq or U
c Probabilistic methods where the CPT readings have been calibrated from lab tests on
recovered soil samples (eg Tumay et al 2013)
A121 Approximate Rules of Thumb
The approximate rules of thumb provide simple guidelines for soil type and suggest that sands are
identified when qt gt 5 MPa and u2 asymp u0 whereas intact clays are prevalent when qt lt 5 MPa and u2 gt u0
(Mayne et al 2002) The magnitude of porewater pressures reflects the consistency of the intact clay
such that soft (u2 asymp 2middotu0) firm (u2 asymp 4middotu0) stiff (u2 asymp 8middotu0) and hard (u2 asymp 20middotu0) However for fissured
overconsolidated clays the measured porewater pressures are less than hydrostatic in fact often
negative u2 lt 0 On land negative porewater pressure readings at the shoulder filter element of the
piezocone should be greater than -1 bar ie u2 gt -100 kPa (-147 psi) Also for clean sands the friction
ratio (FR = 100middotfsqt lt 1) while for insensitive clays (FR gt 4)
A-9
Figure A5 presents a 36-m deep piezocone sounding from the MnDOT Wakota Bridge site which crosses
the Mississippi River at I-494 (Dasenbrock 2006) The qt and u2 readings have been annotated using the
aforementioned rules of thumb indicating the predominance of sands at this site As noted there are
five interbedded clay layers evident at depths of 1 m 3-4 m 7 - 12 m 17 m and 22-27 m
Figure A5 Interpretation of soil types from CPT data taken from the Wakota Bridge on I-494 using
approximate rules of thumb
A122 Soil Behavioral Type Charts
The most common approach for identification of soil types is based on soil behavioral charts
Reviews of several chart methods are provided elsewhere (Kulhawy amp Mayne 1990 Fellenius amp
Eslami 2000 Shahri et al 2015) Current popularity favors the 9-zone classification system to
evaluate soil behavioral type (SBT) that uses normalized piezocone readings (Robertson 1990
1991 Lunne et al 1997)
A-10
(119902119905 minus 120590119907119900) (a) normalized tip resistance 119876 = A11 prime 120590119907119900
100∙119891119904 (b) normalized sleeve friction 119865() = A12
(119902119905 minus 120590119907119900)
(1199062 minus 1199060) (c) normalized porewater pressure 119861119902 = A13
(119902119905 minus 120590119907119900)
While the Q-F-Bq classification is a three-dimensional plot it is often presented as a set of two-dimensional
plots specifically Q vs F and Q vs Bq as shown in Figure A6 Data are grouped according to layers and
can be superimposed to identify their association with the nine soil behavioral types All indirect CPT soil
classification approaches should be cross-checked and verified for a particular geologic setting and local
geotechnical conditions before routine use in practice
Figure A6 Nine-zone soil behavioral type charts (a) normalized cone resistance vs normalized
sleeve friction (b) normalized cone resistance vs normalized porewater pressure parameters
(adapted after Robertson 1991 Lunne et al 1997)
The development of a CPT material index (Ic) has been found advantageous in the initial screening of soil
types and helps to organize the sounding into their respective SBT zones of similar soil response In this
case the CPT index is found from (Robertson 2009)
A-11
A2 22 )log221()log473( rtnc FQI
where Qtn = modfied stress-normalized net cone resistance given by
(119902119905 minus 120590119907119900)120590119886119905119898 119876 = A31 prime (120590119907119900120590119886119905119898)119899
where σatm = reference stress equal to atmospheric pressure (1 atm asymp 1 bar = 100 kPa asymp 1 tsf asymp 147 psi)
The exponent n is soil-type dependent n = 1 (clays) n asymp 075 (silts) and n asymp 05 (sands) If units of bars are
used then Qtn (units of bars) is simply
(119902119905 minus 120590119907119900) 119876 = 119899 A32
prime ) (120590119907119900
The operational value of exponent n is found by iteration Initially an exponent n = 1 is used to calculate
the starting value of Ic (ie Qtn = Q) and then the exponent is upgraded to
prime 120590119907119900 119899 = 0381 ∙ 119868119888 + 005 ( ) minus 015 le 10 A4 120590119886119905119898
Then the index Ic is recalculated Iteration converges quickly and generally only 3 cycles are needed to
secure the operational Ic at each depth The soil zones and associated Ic values are detailed in Figure A7
The sensitive soils of zone 1 can be screened using the following expression
A-12
Zone 1 119876119905119899 lt 12119890119909119901( minus14 ∙ 119865119903)
A5
If any soil layers are found within zone 1 (sensitive soils) then caution should be exercised as these
structured clays are prone to instability collapses and difficulties in construction and performance
Another indicator of zone 1 sensitive soils is when the normalized porewater parameter Bq gt 08 When
these criteria are evident the geoengineer should consult with senior geotechnical personnel or the chief
engineer for guidance
Figure A7 Delineation of soil behavioral type zones using CPT material index Ic
A-13
Stiff overconsolidated clayey sands and sandy clays soils of zone 8 (15 lt Fr lt 45) and zone 9 (Fr gt
45) can be identified from the following criterion
Zones 8 and 9 0020)1(00030)1(0050
12
rr
tnFF
Q A6
The remaining soil types are identified by the CPT material index Zone 2 (organic clayey soils Ic ge 360)
Zone 3 (clays to silty clays 295 le Ic lt 360) Zone 4 (silt mixtures 260 le Ic lt 295) Zone 5 (sand mixtures
205 le Ic lt 260) Zone 6 (clean sands 131 le Ic lt 205) and Zone 7 (gravelly to dense sands Ic le 131)
The red dashed line at Ic = 260 represents an approximate boundary separating drained (Ic lt 260) from
undrained behavior (Ic gt 260)
Once the specific zone has been assigned to a layer a visual representation can be made to show
either the zone number or colorization so that the predominant layers and soil types can be
realized
A123 Probabilistic Soil Types from CPT
The inference of soil type has been addressed using probabilistic methods as discussed by Abu-
Farshakh et al (2008) and Tumay et al (2013) Here the CPT readings can be post-processed to
provide the percentage of sand silt and clay components with depth The output is consistent
and compatible with the current MnDOT soil classification chart (Figure A8) which is similar in
content to that used by the USDA
A-14
Figure A8 Chart for soil classification system by MnDOT
A13 Identifying Soft Organic Soils
The identification of soil type using cone penetration tests (CPT) is usually done on the basis of empirical
soil behavioral type (SBT) charts that use the measured readings (qt fs andor u2) While there at least
25 different sets of charts available (ie Hegazy 1998 Eslami et al 2017 Valsson 2016) one of the most
widely used and popular is the 9-zone SBT system that uses three normalized cone readings Qtn Fr and
Bq (Robertson amp Cabal 2007 Lunne et al 1997) The system is denoted as SBTn and plotted on charts in
terms of Q vs F and Q vs Bq as shown in Figure A9
A-15
1
10
100
1000
01 1 10
Norm
aliz
ed
Con
e T
ip R
esis
tan
ce
Q
Normalized Friction Fr = 100 fst(qt - svo) ()
9-Zone Soil Behavioral Type Chart
Gravelly Sands(zone 7)
Sands
(zone 6)
Sandy Mixtures(zone 5)
Silt Mix(zone 4)
Clays
(zone 3)
Organic Soils(zone 2)
Sensitive Clays
and Silts(zone 1)
Very stiff OC clayto silt (zone 9)
Very stiffOC sandto clayey
sand(zone 8)
1
10
100
1000
-06 -04 -02 00 02 04 06 08 10 12 14
No
rmal
ized
Co
ne
Tip
Res
ista
nce
Q
Normalized Porewater Bq = Du2(qt-svo)
Clays(zone 3)
Sensitive(zone 1)Organic
(zone 2)
Gravelly Sands(zone 7)
Sands(zone 6)
Figure A9 CPT charts for SBTn system (a) Q versus F (b) Q versus Bq
Of particular concern in geotechnical site characterization is the proper identification of soft
organic soils such as organic clays and silts (OL and OH) and peats (Pt) These soils often cause
difficulties and problems during construction and performance of civil engineering works because
of bio-degradation long-term creep excessive settlements andor other issues
The SBTn system has a category labelled Zone 2 recognizes organic soils However several recent
studies have noted that data from CPTu soundings do not always fall within the bounds
delineated using Zone 2 even though laboratory tests and field inspections clearly note the
presence of organic soils Lab tests to classify organic soils includes loss on ignition (LOI gt 5)
and two pairs of Atterberg limits testings per ASTM standards as well as the visual-manual
identification of strong organic odor and smell and dark color notably black and dark gray
Results by Mlynarek et al (2014) show that organic soils in Poland are not adequately identified by the
Q-F chart as seen in Figure A10 Moreover studies by Zawrzykraj et al (2017) found that neither the Q-
A-16
F and Q-Bq plots worked well to recognize organic soils and peats in Poland Nejaim et al (2016) found
that Brazilian soft organic clays were misclassified as Zone 3 (clays) rather than Zone 2 (organic clays)
Similarly a recent study on CPTu data from organic clays located at 24 sites in the USA Sweden Mexico
Brazil and Australia have been compiled (Agaiby 2018) These data are shown to generally avoid falling
with the Zone 2 boundaries in either Q-F or Q-Bq charts thus are not recognized during these soundings
as indicated by Figure A11 The exception in this case is the Mexico City clay which is correctly identified
however the other 23 sites are generally not properly recognized and categorized Additional findings by
Missiaen et al (2015) found that the CPTu results in Belgian soils also did not identify organic clays and
peats in the non-normalized version of these charts that uses Qt versus friction ratio (Rf = 100middotfsqt) as
detailed by Robertson amp Cabal (2007)
Figure A10 Soil behavioral chart with superimposed CPTu data from Polish organic soils
(from Mlynarek et al 2014)
A-17
Figure A11 Paired SBTn charts with CPTu data from 24 organic clay sites indicating non-compliance
with Zone 2 (organic soils)
(compiled by Agaiby 2018)
A14 Simplified Yield Stress Evaluation in Clays by CPTu
From the development of a hybrid formulation of spherical cavity expansion (SCE) theory with critical-
state soil mechanics (CSSM) the effective preconsolidation stress or yield stress (σp) of clays can be
expressed in terms of three piezocone parameters (a) net cone tip resistance qnet = qt - σvo (b)
measured excess porewater pressure Δu2 = u2 - u0 and (c) effective cone resistance qE = qt - u2 Details
on the SCE-CSSM solutions are given elsewhere (Mayne 1991 Mayne 2017 Agaiby amp Mayne 2018)
A-18
For normal and well-behaved clays which include inorganic fine-grained soils such as clays and silts
of low sensitivity a set of simple relationships can be derived By adopting characteristic default values
for effective friction angle ϕ = 30deg and rigidity index (IR = 100) linear expressions for the effective
preconsolidation stress are obtained
prime 120590119901 = A7 033 ∙ 119902119899119890119905
prime 120590119901 = 053 ∙ ∆1199062 A8
prime 120590119901 = 060 ∙ 119902119864 A9
A141 Application to soft Chicago clay
An example of a common geomaterial in this category includes the infamous soft Chicago clays
that were deposited in a glacio-lacustrine environment (Cho amp Finno 2010) The soft clay has a
mean water content of 20 liquid limit of 38 and plasticity index PI= 12 Results of CPTu tests
conducted at the national geotechnical experimentation site (NGES) at Northwestern University
are shown in Figure A12 (Ouyang amp Mayne 2017) As seen in the profile a thin soft clay resides
between depths of 9 to 105 m with a thicker soft clay layer in the range of 12 to 18 m
A-19
0
2
4
6
8
10
12
14
16
18
20
00 02 04 06 08 10 12D
ep
th (
m)
CPTu qt and u2 (kPa)
sand (SP)
Blodgett
Park Ridge
Deerfieldsoft clay
ClayLayerof Study
Northwestern UnivNational Test Site
Figure A12 Results of CPTu sounding at the NGES at Northwestern University Illinois
Post-processing of the CPTu data using Equations A7 A8 and A9 show good agreement in evaluating the
profile of effective yield stress in this clay layer compared with results from one-dimensional
consolidation tests made on undisturbed samples taken at the site
A-20
10
11
12
13
14
15
16
17
18
19
20
0 100 200 300 400 500 600 700 800 900 1000
De
pth
(m
)Effective Yield Stress sp (kPa)
Consols
033 qnet
053 Du2
060 qE
svo
Figure A13 Evaluation of effective yield stresses at NWU using consolidation tests and CPTu
A142 Application to soft Bay Mud California
Another example of a well-behaved fine-grained soil is the San Francisco Bay Mud which is a soft clay
deposit in northern California Results of a representative CPTu are shown in Figure A14 and indicate the
soil profile at a site in eastern San Francisco (Hunt et al 2002) For this site index values include 70 lt
wn lt 90 60 lt LL lt 80 and 35 lt PI lt 45 Post-processing the CPTu data using Equations A7 A8 and
A9 show good agreement with each other as well as with the results of consolidation tests as seen in
A-21
Figure A15 The consolidation tests were performed using a CRS device as reported by Pestana et al
(2002)
San Francisco Bay Mud (Pestana et al 2002)
0
5
10
15
20
25
30
35
40
0 1000 2000 3000
De
pth
(m
ete
rs)
Cone Resistance qt (kPa)
0
5
10
15
20
25
30
35
40
0 10 20 30 40 50 60
Sleeve Friction fs (kPa)
0
5
10
15
20
25
30
35
40
0 500 1000 1500
Porewater u2 (kPa)
Miscellaneous Fill
Young Bay Mud
Young Bay Mud
Clayey Sand
Miscellaneous Fill
Young Bay Mud
Young Bay Mud
Figure A14 Representative CPTu in San Francisco Bay Mud (data from Hunt et al 2002)
A-22
San Francisco Bay Mud (Pestana et al 2002)
0
5
10
15
20
25
30
0 100 200 300 400 500 600
De
pth
(m
ete
rs)
Preconsolidation Stress sp (kPa)
svo
033 qnet
053 Du2
060 qE
CRS
svo
Figure A15 Evaluation of effective yield stresses in Bay Mud from consolidation tests and CPTu
A15 CPTu Screening of Soft Organic Clays
For soft organic clays the Equations A7 A8 and A8 will not apply thus can serve as a means to
screen such geomaterials from the soil profile Two examples of CPTu soundings in soft organic
clays are presented to illustrate the approach Additional case records and documentation of
CPTu data in organic clays are given in Agaiby (2018)
A151 Soft organic alluvial soils in Washington DC
Results of in-situ tests including CPTu soundings in soft organic clayey silts along the Potomac River at the
Anacostia Naval Air Station are presented by Mayne (1987) Here the soft soils are categorized as organic
clayey silts (OH) per the Unified Soils Classification Systems (USCS) with mean values (and plus and minus
A-23
one standard deviations) of wn = 68 plusmn 16 LL = 83 plusmn 25 and PI = 37 plusmn 17 A representative
piezocone sounding is depicted in Figure A16 For the post-processing of CPTu data the use of the Q-F
and Q-Bq graphs do not find any points within the Zone 2 category for organic soils When the data are
evaluated to determine the profile of effective yield stress the Equations A7 A8 and A9 do not agree as
shown by Figure A17
0
5
10
15
20
25
0 2000 4000 6000
Dep
th (
met
ers)
Cone Resistance qt (kPa)
0
5
10
15
20
25
0 20 40 60 80 100
Sleeve Friction fs (kPa)
0
5
10
15
20
25
0 200 400 600
Pore Pressure u2 (kPa)
Figure A16 Representative CPTu in soft organic clay along the Potomac River DC
A-24
0
5
10
15
20
25
0 3 6 9D
epth
(m
eter
s)Soil Behaviour Type Zone
Q and Fchart
0
5
10
15
20
25
0 100 200 300 400 500 600
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
033qnet
053 Du2
06 qE
Oed
Figure A17 Post-processing of CPTu data from Anacostia-Bolling (a) SBTn zones using CPT material
index Ic (b) CPTu screening equations for yield stress
Note that the hierarchy of the results from Equations A7 A8 and A9 show that the CPTu
evaluations for preconsolidation stress in soft organic clays indicates
[A10] σp = 053 Δu2 lt 033 qnet lt 060 qE
A152 Soft organic peaty clay in Saint Paul MN
Results of CPTu tests were collected during a training exercise underneath I-35E just northeast
of Saint Paul MN in 2007 All three MnDOT CPT rigs available at the time were used to obtain
in-situ test data at the site The site is well known as having soft organic soils and samples were
obtained from three borings made at the site (Boring ID T523 T542 and T518) Boring logs
A-25
indicate the presence of highly organic silt loams with marl and spongy organic clayey silts
From 12 samples the natural water contents ranged from 30 to 191 with a mean of 112 plusmn
58
A representative piezocone sounding from the site (MnDOT CPTu ID F22Y0703C) is used for
this illustration and is presented as Figure A18 Post-processing these data using the SBTn
algorithms and CPT material index show that the soils classify primarily as Zone 3 (clays) and
Zone 4 (silty mix) thus do not identify the geomaterials properly as Zone 2 (organic soils)
0
5
10
15
0 500 1000 1500 2000
Dep
th (
met
ers)
Cone Tip Resistance
qt (kPa)
0
5
10
15
0 10 20 30 40 50 60
Sleeve Friction
fs (kPa)
0
5
10
15
-100 0 100 200 300 400
Porewater Pressure
u2 (kPa)
Figure A18 MnDOT CPTu sounding at the I-35E test site near St Paul Minnesota
A-26
1
10
100
1000
01 1 10
Norm
aliz
ed
Con
e T
ip R
esis
tan
ce
Q
Normalized Friction Fr = 100 fst(qt - svo) ()
9-Zone Soil Behavioral Type Chart
St PaulGravelly Sands(zone 7)
Sands
(zone 6)
Sandy Mixtures(zone 5)
Silt Mix(zone 4)
Clays
(zone 3)
Organic Soils(zone 2)
Sensitive Clays
and Silts(zone 1)
Very stiff OC clayto silt (zone 9)
Very stiffOC sandto clayey
sand(zone 8)
1
10
100
1000
-06 -04 -02 00 02 04 06 08 10 12 14
No
rmal
ized
Co
ne
Tip
Res
ista
nce
Q
Normalized Porewater Bq = Du2(qt-svo)
Clays(zone 3)
Sensitive(zone 1)Organic
(zone 2)
Gravelly Sands(zone 7)
Sands(zone 6)
Figure A19 Post-processing St Paul sounding using the 9-zone SBTn classification system
Using the recommended approach Figure A20 shows the CPT-evaluated yield stress profiles from
equations A7 A8 and A9 Here again the three estimates of σp do not agree but show the same
hierarchy as detailed in Equation A10
A-27
0
5
10
15
20
0 50 100 150 200 250 300 350 400
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
06 qeff svo
033qnet Consols
054 Du2
Figure A20 CPTu screening of soft organic soils at St Paul test site
A16 CPTu Evaluations of Yield Stress in Soft Organic Soils
Once the presence of soft organic clays and silts is identified using the aforementioned CPT
screening algorithms the evaluation of effective yield stress is conducted using the procedure
outlined in Chapter 2 of this MnDOT manual For soft organic soils an exponent of m = 090 is
recommended (Mayne et al 2009) Therefore the preconsolidation stress is obtained from
(units of kPa)
prime = A11 120590119901 033 ∙ (119902119899119890119905)090
The algorithm can be expressed in dimensionless form by
prime = A12 120590119901 033 ∙ (119902119905 minus 120590119907119900)119898prime ∙ (120590119886119905119898100)1minus119898prime
A-28
where σatm = reference stress (= 100 kPa = 147 psi) and m = 090 for soft organic silts and clays
The approach is applied to the two former example case studies in Figure A21 (Washington DC)
and Figure A22 (St Paul MN) respectively showing good agreement with benchmark results
taken from laboratory consolidation tests on undisturbed samples of these sites in both cases
0
5
10
15
20
25
0 100 200 300 400 500 600
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
033qnet^090
Oed
Figure A21 Profiles of yield stress from consolidation tests and CPTu method for organic clayey silts at
Anacostia-Bolling site in Washington DC
A-29
0
5
10
15
20
0 20 40 60 80 100 120 140 160
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
CPT
Consols
120590119901prime = 033 119902119899119890119905
090 in kPa
Figure A22 Profiles of yield stress from consolidation tests and CPTu method for organic clayey silts at
MnDOT test site near Saint Paul Minnesota
A17 Soil Unit Weight
As shown by Figure A23 total soil unit weight (t) can be estimated from the CPT sleeve friction resistance
(Mayne 2014) The equation can be expressed by
120574119905 = 120574119908 ∙ [122 + 015 ∙ 119897119899 (100 lowast 119891119904120590119886119905119898 + 001)] A13
where w = unit weight of water and satm = atmospheric pressure
A-30
Figure A23 Evaluation of soil unit weight from CPT sleeve friction
A18 Effective Stress Friction Angle
Sands
The effective stress friction angle () is one of the most important soil properties as it governs the strength
of geomaterials as well as affects soil-pile interface and pile side friction While an effective cohesion
intercept (c) can also be considered this is usually reserved for cemented or bonded geomaterials or
unsaturated soils and may lose its magnitude with time ageing or with prolonged environmental
changes
For clean quartz to silica sands where porewater pressures are essentially hydrostatic (Bq = 0) the
following expression has been calibrated with triaxial compression test results from undisturbed sand
samples and normalized cone resistances as presented in Figure A24 (Mayne 2007 2014)
A-31
120601prime(119889119890119892) = 176deg + 110deg 119897119900119892 (1199021199051) A14
Figure A24 Evaluation of effective stress friction angle in quartz-silica sands from CPT
Note Relationship applies to drained soil behavior when Ic lt 26 andor Bq lt 01
where qt1 is an earlier form of stress normalized cone tip resistance for sands given by
(119902119905 120590119886119905119898) 1199021199051 = A151 05 prime (120590119907119900120590119886119905119898)
A-32
In terms of units of bars this is expressed simply as (in units of bars)
119902119905 1199021199051 = 05 A152 prime ) (120590119907119900
Recently Robertson amp Cabal (2015) recommended the use of the modified form of normalized cone
resistance (Qtn) in Equation A10
120601prime (119889119890119892) = 176deg + 110deg 119897119900119892 (119876119905119899) A16
In sands qnet asymp qt since the overburden term is small relative to the cone tip resistance Figure A25 shows
that the two normalizations give very comparable results
Figure A25 Comparable values from qt1 and Qtn normalization schemes for CPTs in sands
Clays and Silts
A-33
In the case of soft to firm intact clays and silty clays the effective friction angles are determined from the
normalized cone resistance and porewater pressure parameters (Senneset et al 1989 Mayne 2016) as
shown in Figure A26 The exact solution when the angle of plastification = 0 is given as
qBQ
)tan1(tan61
1)tanexp()245(tan 2
A17
Approximate NTH Solution for from CPTu
qBQ
)tan1(tan61
1)tanexp()245(tan2
Approx asymp 295deg∙Bq0121∙[0256 + 0336∙Bq + log Q ]
1
10
100
20 25 30 35 40 45
Co
ne
Res
ista
nce
Nu
mb
er Q
(degrees)
Bq
010204060810
c = 0 = 0deg
Theory (dots)
Approximation (lines)
Theory
Figure A26 Evaluation of effective stress friction angle in clays and silts from CPTu results
Note Relationship applies to undrained soil behavior when Ic ge 26 andor Bq ge 01
which can be approximately inverted into the form (Mayne 2007)
A-34
0121 120601prime = 295deg ∙ 119861119902 ∙ [ 0256 + 0336 ∙ 119861119902 + log(119876)] A18
This algorithm is specifically applicable for the following ranges of porewater pressure parameter (01 le
Bq lt 1) and effective stress friction angles (20deg le lt 45deg)
A19 Stress History
The stress history can be characterized by an apparent yield stress or preconsolidation stress (sp) as well
as by its normalized and dimensionless form YSR = spsvo = yield stress ratio The YSR is in effect the
same as overconsolidation ratio (OCR) however now generalized to accommodate mechanisms of
preconsolidation that occur beyond just erosion glaciation and removal of overburden stresses but also
due to ageing desiccation repeated cycles of wetting-drying bonding repeated freeze-thaw cycles
groundwater changes and other factors
A-35
A generalized approach for evaluating the yield stress or preconsolidation in soils using net cone
resistance has been formulated as presented in Figure A27 (Mayne 2015) Yield stress in soils from CPT
10
100
1000
10000
10 100 1000 10000 100000
Yie
ld S
tre
ss
s
y
(kP
a)
Net Cone Resistance qt - svo (kPa)
Blessington
General Trend
sy = 033(qt-svo)m
Fissured Clays m = 11
Intact clays m = 10 Sensitive clays m = 09
Silt Mixtures m = 085Silty Sands m = 080
Clean Sands m = 072
m = 11 10 09
085
080
072
Units kPa
Fissured Clays
Figure A27 Evaluation of yield stress or preconsolidation stress in soils from CPT
The algorithm can be expressed in dimensionless form by
1minus119898prime prime 120590119886119905119898 120590119901 = 033 ∙ (119902119905 minus 120590119907119900)119898prime
( ) A19 100
where m = exponent depends on soil type m = 1 (intact clays) 085 (silts) 080 (sandy silts to silty sands)
and 072 (sands) For fissured clays the exponent m may be 11 or higher depending upon the age of the
A-36
formation and degree of jointing and discontinuities Note that fissured clays can be identified when Ic lt
26 and Bq lt 01
The exponent m has also been calibrated with CPT material index as presented in Figure A28
An algorithm to express this relationship for non-fissured soils is given by
25)652(1
2801
cIm
A20
This allows the post-processing of CPT to automatically choose the appropriate exponent m for
a layer by layer analysis
Figure A28 Yield stress exponent m in terms of CPT material index Ic
A-37
A110 Lateral Stress Coefficient
The horizontal geostatic state of stress is represented by the lateral stress coefficient K0 = shosvo
commonly referred to as the at-rest condition The magnitude of K0 for soils that have been loaded and
unloaded can be approximately estimated from
119870119900 = (1 minus 119904119894119899120601prime) ∙ 119874119862119877119904119894119899120601prime A21
Data compiled from in-situ K0 measurements using self-boring pressuremeter tests (SBPMT) total stress
cells (TSC) or push-in spade cells andor laboratory methods (instrumented consolidometers triaxials
andor suction measurements) on a variety of clays silts and sands have been compiled and reported by
Ku amp Mayne (2015) as shown in Figure A29 verifying Equation A15 as a means for evaluating K0 in soils
Figure A29 Relationship between lateral stress coefficient K0 and YSR or OCR in soils (Ku amp Mayne
2015)
A-38
A111 Undrained Shear Strength
The loading of soils can occur under conditions of being fully drained (Du = 0) partially drained or full
undrained (DVV0) where Du = excess porewater pressures (above hydrostatic) and DVV0 = volumetric
strain Further details on specific stress paths are best explained in terms of critical state soil mechanics
or CSSM (Mayne et al 2009 Holtz Kovacs amp Sheahan 2011) The prevailing drainage conditions depend
upon the rate of loading and permeability characteristics of the soil Normally in sands that are pervious
and exhibit high permeability a drained response occurs Exception to this may occur in loose sands
during fast earthquake loading resulting in soil liquefaction In clays that exhibit low permeability a fast
rate of loading will result in undrained loading at constant volume This in fact is a temporary and
transient condition often termed short term loading In the long term eventually porewater pressures
will dissipate (albeit slowly) and a drained response will prevail thus termed long term loading
The overall soil strength is controlled by the effective stress strength envelope Most commonly this is
represented by a simple linear relationship termed the Mohr-Coulomb criterion where the maximum
shear stress (max called the shear strength) is given as
= 119888prime + 120590prime ∙ 119905119886119899 120601prime A22 120591119898119886119909
where c = effective cohesion intercept s = effective normal stress and = effective stress friction angle
As a starting point values of c = 0 and = 30deg can be adopted for all soil types at least until the specific
soil type geologic formation and results from high-quality laboratory or field test data are available
Peak Undrained Shear Strength from Stress History
Using simplified critical state soil mechanics the peak undrained shear strength (max = su) can be
evaluated from the effective stress strength envelope (c = 0 effective friction angle ) and stress history
(ie OCR) in the form
prime DSS 119904119906 = (119904119894119899120601prime2) ∙ (119874119862119877)Λ ∙ 120590119907119900 A23
A-39
which corresponds to a simple shear mode Figure A30 presents the aforementioned relationship
together with data from 17 different clays tested in simple shear
Figure A30 Normalized undrained shear strength from simple shear tests on various clays
versus YSR or OCR
For the triaxial compression (TC) mode the equation would give slightly higher strengths that are
calculated from
prime TC 119904119906 = (1198721198882) ∙ (1198741198621198772)Λ ∙ 120590119907119900 A24
where Mc = (6 middot sin)(3-sin)
Peak Undrained Shear Strength from CPTu
A more direct approach to assessing undrained shear strength is via bearing capacity theory
whereby
A-40
A25 kt
votu
N
qs
s
where Nkt = bearing factor that depends on mode of shearing sensitivity of the clay degree of
overconsolidation and other factors For soft offshore clays the back figured Nkt from data collected at
14 well-documented sites (Low et al 2010) determined mean values based on mode of shearing Nkt =
119 (triaxial compression) Nkt = 136 (simple shear) and Nkt = 133 (vane)
A recent study of 51 clays that were tested by both field piezocone and laboratory CAUC triaxial tests
showed that essentially Nkt = 12 for intact clays and clayey silts of low to medium sensitivity (Mayne
Peuchen amp Baltoukas 2015) Figure A31 shows the slightly different trends for offshore versus onshore
clays For sensitive clays a lower value of Nkt = 10 would be appropriate and for fissured clays a higher
value of Nkt would be in the range of 20 to 30
A-41
A-42
Figure A31 Database from 51 clays with CPT qnet versus lab measured triaxial shear strength
(Mayne Peuchen amp Baltoukas 2015)
For soft to firm clays an alternate means to evaluate undrained shear strength is via the excess porewater
pressures ( u = u2 - u0) from the following
u
uN
us
D
D 2
A26
where N u = porewater bearing factor also dependent on the aforementioned factors The study by
Low et al (2010) found N u = 59 (triaxial compression) N u = 69 (simple shear) and N u = 71 (vane
shear)
A third alternate is found from the effective cone resistance (qE = qt - u2) whereby
kE
tu
N
uqs 2
A27
where NkE is a bearing term for this approach Mayne et al (2015) indicated a mean value of NkE = 8 for
soft-firm intact clays
Remolded Undrained Shear Strength from CPT
The remolded undrained shear strength (sur) is obtained from either field vane tests lab mini-vane shear
tests or lab fall cone devices This affords the evaluation of the clay sensitivity (St) which is defined as the
ratio of peak to remolded strengths at a given water content
119904119906(119901119890119886119896) 119878119905 = frasl A28 119904119906119903
For the CPT the sleeve friction has been noted to give values that are comparable to the
remolded undrained shear strength (Powell amp Lunne 2015) Therefore
asymp 119891119904 A29 119904119906119903
A112 Ground Stiffness and Soil Moduli
The stiffness of the ground can be represented by several geoparameters including the compressibility
indices (Cr Cc Cs) spring constants (kv) and moduli Regarding the latter there are several moduli that
are used in geotechnical engineering including shear modulus (G and Gu) Youngs modulus (E and Eu)
constrained modulus (D) bulk modulus (K) resilient modulus (MR) and subgrade reaction modulus (ks)
A-43
Soil Modulus
The definition of modulus is taken as E = DsD with Figure A32 showing a typical deviator stress (q = s1
- s3) versus axial strain (a) curve Theoretical interrelationships between the elastic moduli G E D and
K have a dependence on the Poissons ratio v The value of Poissons ratio can be taken as vu = 05 for
undrained loading (ie constant volume) while for drained loading which is accompanied by volumetric
strains a value of v = 02 may be used If we adopt the reference modulus as E then the
interrelationships with the other elastic moduli are given by
a Reference Stiffness 119864prime = drained Youngs modulus A30
119864prime
b Shear Modulus 119866prime = A31 [2(1+120584prime)]
119864prime (1minus120584prime) c Constrained Modulus 119863prime = A32
[(1+120584prime)(1minus2120584prime)]
119864prime
d Bulk Modulus 119870prime = A33 [3middot(1minus2120584prime)]
A-44
Figure A32 Representative stress-strain-strength curve for soils in triaxial compression
For the extreme case when = 0 in fact D = E When = 02 the two moduli are only 10
different D = 11middotE thus the constrained modulus and drained Youngs modulus are often
considered somewhat interchangeable
The resilient modulus (MR) applies to pavement analysis and design most commonly measured by cyclic
triaxial testing under repeated load applications In fact MR is a special case of Youngs modulus that
relates to the small strain stiffness measured in the nondestructive range but has developed permanent
plastic strains after many cycles of loading (Brown 1996 Dehler amp Labuz 2007)
The subgrade reaction modulus is actually a combined soil-structural parameter as its value
depends on the ground stiffness and the size of the loaded element The subgrade modulus is
defined as ks = q where q = applied stress and = measured deflection In terms of elasticity
solutions the deflection of a flexible circle of diameter d is given by = qmiddotdmiddot(1-2)E Therefore
119864prime
119896119904 = A34 [119889middot(1minus1205842)]
which has units of kNm3 or pcf
Table A2 summarizes several selected studies towards the approximate evaluation of these moduli
obtained directly from measured CPT data Note however that the results from in-situ penetrometer
tests generally represent a peak strength as indicated in Figure A32 Thus the measured cone tip
resistance (qt) reflects the top of the stress-strain curve either the undrained strength in clays (su) or the
effective friction angle () of sands or even a strength intermediate between these two values Thus
A-45
Source Soils Studied Reference Modulus
Findings
Kulhawy amp Mayne 1990
8 stiff OC clays Lab constrained modulus
D asymp 825 middot (qt - svo)
Mohammed et al (2000)
Resilient modulus
Abu-Farsakh 2004
7 Louisiana sites
Lab constrained modulus
D asymp 358 middot (qt - svo)
Mayne (2007b)
All soil types Constrained moduli from lab consolidation
First-order estimate D asymp middot(qt - svo)
Robertson (2009)
All soil types Constrained modulus from consolidation
D = m middot (qt - svo)
when Ic gt 22
1 m = Qtn when Qtn le 14
2 m = 14 when Qtn gt 14
when Ic lt 22 then
= 003 middot 10055 Ic + 168 m
Liu et al (2016)
16 clay sites in China
Resilient modulus MR = (146qt 053 + 1355 fs
14 + 236)244
(all in MPa)
Casey et al (2016)
8 clays Triaxial tests for undrained Youngs modulus
Eu(svc)07 = 316 - 23 middot LL
where Eu and svc (MPa)
and LL = liquid limit ()
qt is probably not the best means to obtain the slope of the stress-strain curve Instead the SCPTu that
provides the initial tangent shear modulus (Gmax) from the shear wave velocity (Vs) is a better choice
Table A2 Selected Modulus Relationships with Cone Penetration Test Measurements
A-46
Nonlinear Modulus
The small-strain shear modulus (Gmax or G0) represents the initial stiffness of all soils and rocks It is the
beginning of all stress-strain-strength curves for geomaterials and obtained from elasticity theory
119866119898119886119909 = 1198660 = 120588119905 middot 1198811199042 A35
where Vs = shear wave velocity t = tga = total soil mass density t = total soil unit weight and ga =
gravitational acceleration constant (98 ms2)
From a general viewpoint on stiffness the shear modulus G of soil is defined as the slope of shear stress
versus shear strain G = DDs for a tangent definition and by G = s as a secant definition The shear
modulus is related to its associated Youngs modulus E = 2G(1+v) Both moduli are in fact highly
nonlinear ranging from a maximum value at the small-strain stiffness (Gmax = G0 = t middotVs2 where t =
total soil mass density and Vs = shear wave velocity) to intermediate G values at medium strains (asymp 1)
to low values at peak strength As such a variety of algorithms and formulae have been developed to
represent either a partial range or the full stress-strain-strength behavior of soils over a range of
interests (eg Mayne 2005) In these formulations a variable number of input parameters may be
required in order to produce a stress-strain-strength curve
A modified hyperbolic form suggested by Fahey amp Carter (1993) has favorable attributes in that the
modulus reduction factor can be established with only a single variable This allows the initial stiffness
to the be small-strain stiffness (G0) that is reduced in terms of level of mobilized strength eg max =
qqmax = 1FS which is simply the reciprocal of the calculated factor of safety (FS) The magnitude of
secant shear modulus (G) corresponding to the particular level of loading is given by
119866 = 119872119877119865 middot 1198660 = (1198661198660) middot 1198660 A36
A-47
where the MRF = modulus reduction factor determined as
1 minus (120591 )119892 119872119877119865 = (1198661198660) = frasl A37 120591119898119886119909
and g = empirical exponent term Figure A33 shows the relationships for normalized shear stress ()
versus shear strain (s) and corresponding normalized secant shear modulus (G = s) with shear strain
over a range of exponent values 01 le g le 10 For a discussion of tangent G please see Fahey amp Carter
(1993)
Using laboratory data from resonant column-torsional shear tests and triaxial specimens with local
strain measurements for Gmax reference values modulus reduction curves for a selection of sands and
clays are presented in Figure A34 The data indicate that the exponent g falls within the range 02 lt g lt
05 for many soils tested drained and undrained A mean value of g = 03 is recommended for
preliminary site investigations and designs until additional information can be obtained
The value of max is the shear strength of the soil generally taken as either (1) drained (c = 0) max =
svo middot tan or (2) undrained where max = su = undrained shear strength as discussed earlier Thus each
shear stress () can be associated with its shear modulus (G) and the relevant shear strain is found from
s = G This allows for generation of nonlinear stress-strain-strength curves at all depths from SCPTu
data in clays sands and mixed soil types
A-48
Modified Hyperbola (Fahey amp Carter 1993 Canadian Geotech J) Coefficient Term f = 1
``
0
01
02
03
04
05
06
07
08
09
1
000001 00001 0001 001 01 1
No
rmal
ized
GG
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
0
01
02
03
04
05
06
07
08
09
1
0 01 02 03 04 05 06 07 08 09 1
No
rmali
zed
GG
ma
x
Mobilized Shear Stress max
g =1
07
05
03
02
01
0
01
02
03
04
05
06
07
08
09
1
00001 0001 001 01 1
No
rmali
zed
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
0
01
02
03
04
05
06
07
08
09
1
0 1 2 3 4 5 6 7 8 9 10
No
rmali
zed
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
Reduction Factor
GGmax = 1- (max)g
Figure A33 Normalized modulus (GGmax) and normalized shear stress (max) versus shear strain (s)
for modified hyperbolic relationship of Fahey amp Carter (1993)
A-49
0
01
02
03
04
05
06
07
08
09
1
0 01 02 03 04 05 06 07 08 09 1
Mo
du
lus
Re
du
cti
on
G
Gm
ax
or
EE
ma
x
Mobilized Strength max or qqmax
NC SLB Sand
OC SLB Sand
Hamaoka Sand (e=0625)
Hamaoka Sand
Toyoura Sand e = 067
Toyoura Sand e = 083
Ham River Sand
Ticino Sand
Kentucky Clayey Sand
Kaolin
Fujinomori Clay
Pietrafitta Clay
London Clay
Vallericca Clay
Pisa Clay
Pisa Clay
Pisa Clay
Pisa Clay
Kiyohoro Silty Clay
Thanet Clay
Exp g = 02
Exp g = 03
Exp g = 04
Exp g = 05
MRF = 1 - (qqmax)g
Figure A34 Modulus Reduction Factors (MRF) with mobilized strength of clays and sands
A113 Coefficient of Consolidation
The rate at which foundation and embankment settlements occur as well as the dissipation of excess
porewater pressures is controlled by the coefficient of consolidation (cv) The magnitude of cv is also
required in the design of vertical wick drains that can be installed in soft ground to expedite the time for
consolidation Using the results of CPTu dissipation tests which measure the rate at which the u2 readings
vary with time the in-situ profile of cv can be evaluated Often the piezocone-interpreted values of cv
are validated by comparison with results from laboratory one-dimensional consolidation tests (eg Abu-
Farsakh MY 2004) Alternatively a better method is to cross-check the values with the measured full-
scale performance of instrumented embankments that are constructed over the ground and the recorded
time-rate of consolidation and settlements can provide the best cv for that site (Abu-Farsakh MY et al
2011)
A-50
Monotonic Dissipation Tests
An illustrative dissipation curve from piezocone testing at IDT State Highway 95 at an embankment and
bridge crossing is presented in Figure A35 After the CPTu sounding was halted at a depth of 512 feet
the recorded decay of porewater pressures was observed to be monotonic with time until the dissipations
were ended at 1000 seconds During that time the u2 readings decreased from 115 psi to 47 psi At this
location the depth to the groundwater table is about 16 feet thus the calculated equilibrium water
pressure is 15 psi Commonly a characteristic time for piezo-dissipations is taken at 50 degree of
consolidation although other degrees may be adopted By evaluating the value of u2 at 50 as shown in
Figure A35 the characteristic t50 = 404 s can be obtained
0
20
40
60
80
100
120
1 10 100 1000 10000
Pore
wat
er P
ress
ure
u2
(psi
)
Time (s)
CPTU-02E-1 z = 1560 m
uo (hydrostatic)
u2 (initial) = 115 psi
u2 (final projected) = u0 = 15 psi
u2 at 50 = frac12(115+15) = 65 psi
Idaho DOT State Route 95Dissipation at 512 feet
t50 = 404 seconds
Time to reach50 consolidation
A-51
Figure A35 Piezo-dissipation at Sandpoint Bridge Crossing Idaho for depth z = 512 feet illustrating the
evaluation of characteristic time t50
It is also common to plot normalized excess porewater pressures relative to the measured initial Du2
that is obtained during penetration at the constant rate of 20 mms ie DuDui versus time Figure A36
shows a set of piezo-dissipation records for a bridge and embankment site in southern Louisiana
reported by the LTRC While the porewater pressure axis is arithmetic the time scale is often plotted on
either logarithmic or square root scales In either case the t50 is then found from the measured
dissipation curve when DuDui = 050
A-52
Figure A36 Set of monotonic piezo-dissipations taken at various depths for Courtableau Bridge
Site on Louisana State Highway 103 (Abu-Farsakh et al 2011)
Dilatory Dissipation Curves
In a number of situations a monotonic type dissipation does not occur on the onset but instead a dilatory
type curve is observed Here after stopping the push of the penetrometer the recorded porewater
pressures initially begin to rise and eventually reach a peak value then afterwards follow a phase where
the Du values decrease to hydrostatic (Figure A37) A full solution is available for both cases (Burns amp
Mayne 2002) Dilatory responses are most often associated with overconsolidated clays and silts
although occasionally are observed in fine-grained soils with low OCRs
The selection of the characteristic t50 value may found by using a square root of time plot to find the initial
u2 value by projection of the post-peak dissipatory data back to the ordinate axis as shown by the example
in Figure A38 In this example the initial u2 = 480 kPa and then the same procedures in Figure A35may
apply
A-53
Figure A37 Monotonic and dilatory porewater dissipation responses in soils
(after Burns amp Mayne 1998)
Figure A38 Method for obtaining t50 from dilatory dissipation curves
(Schneider amp Hotstream 2010)
A-54
Interpretation of Dissipation Test Data
There are a number of different solutions available for the interpretation of the coefficient of
consolidation (cv) from the piezocone dissipation curves For monotonic type response the strain path
method (SPM) developed at Oxford University (Teh amp Houlsby 1991) is well-recognized while the Georgia
Tech solution by Burns amp Mayne (2002) is based on spherical cavity expansion and critical state soil
mechanics (SCE-CSSM) and handles both monotonic and dilatory curves For both solutions a simplified
procedure can be recommended that relies on the aforementioned t50 value obtained from the measured
field data as given in Table A3 The units for cv are in cm2s as shown or equivalent
Table A3 Recommended procedures for calculating cv from t50 obtained in dissipation tests
Method Equation for coefficient of
consolidation cv
RemarksNotes Eqn
No
SPM
(Teh amp
Houlsby
1991) 50
2)(2450
t
Iac
Rc
v
t50 = measured time to reach 50
dissipation
IR = Gsu = undrained rigidity index
G = shear modulus
su = undrained shear strength
ac = penetrometer radius (ac = 178
cm for 10-cm2 cone ac = 220 for a
15-cm2 size)
A32
SCE-CSSM
(Burns amp
Mayne 2002) 50
7502 )()(0300
t
Iac Rc
v
A33
A-55
Evaluating rigidity index
Both the SPM and SCE-CSSM solutions require an estimate of the in-situ rigidity index (IR = Gsu) of the
soil If the results of SCPTU are available Krage et al (2014) have derived an expression for IR that
depends upon the small-strain shear modulus net cone tip resistance and effective overburden stress
250750
max50
8111
vonet
Rq
GI
s A38
where consistent units are input for Gmax qnet and svo The value of Gmax is obtained from B29 using the
measured shear wave velocity (Vs) and unit weight evaluated from B7
If the shear wave velocity is not measured it can be estimated from the CPT data A number of
methods have been reviewed for CALTRANS by Wair et al (2012) including one that is applicable
to sands silts and clays of low sensitivity that are inorganic and uncemented
119881 (119898119904) = [101 middot 119897119900119892(119902119905) minus 114]167 middot (100 middot 119891119904119902119905)03 A39 119904
where qt is in units of kPa (Hegazy amp Mayne 1995) An alternative relationship is given by Robertson amp
Cabal (2015)
A114 Hydraulic Conductivity
The hydraulic conductivity is a parameter that expresses the flow characteristics of the soil In
geotechnics it is also called the coefficient of permeability (k) and has units of cms or feetday
Through consolidation theory the hydraulic conductivity relates directly to the coefficient of
consolidation
A-56
A40 D
ck wv
where w = unit weight of water and D = constrained modulus
Therefore one approach to evaluating k can be from the site-specific cv obtained from the
dissipation tests using an estimate of D from one of the relationships given in Table A2
An approach developed for soft normally-consolidated soils that uses t50 directly to assess the magnitude
of k is presented in
Figure A39 (Parez and Fauriel 1988) The mean trendline through the wedges of soil type can be
approximated by
251
50 (sec)251
1)(
tscmkh
A41
A-57
Figure A39 Hydraulic conductivity versus dissipation time for 50 consolidation
(Parez amp Fauriel 1988)
A-58
APPENDIX B
List of Figures
Figure B1 Conventional methods vs direct CPT approach to shallow foundation response B-3
Figure B2 Direct bearing capacity relationship for embedded square and strip footings situated on clean
sands (after Schmertmann 1978) B-6
Figure B3 Direct CPT bearing factor Rk as function of footing width B and embedment depth from finite
element analyses (Tand et al 1995) B-6
Figure B4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio and sand
consistency (Eslaamizaad amp Robertson 1996) B-7
Figure B5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp Salgado (2005) in
terms of footing size relative density and base settlement B-8
Figure B6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and normalized
cone tip resistance (after Eslami amp Gholami 2005) B-8
Figure B7 Direct relationship between ultimate bearing stress in clays and measured cone tip resistance
(Schmertmann 1978) B-10
Figure B8 Direct relationship between ultimate foundation bearing stress in clays and net cone tip
resistance (after Tand et al 1986) B-11
Figure B9 Measured load-displacements for each of the five large footings at TAMU (Briaud amp Gibbens
1994) sand site B-12
Figure B10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU (Briaud amp
Gibbens 1994) footings B-13
Figure B11 Characteristic stress versus square root normalized displacements for TAMU footingsB-15
Figure B12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sandB-16
Figure B13 Summary of sand formation factors with corresponding CPT resistancesB-19
Figure B14 Normalized foundation stresses vs square root of normalized displacement for spread
footings on sand (modified after Mayne et al 2012) B-20
Figure B15 Definitions of capacity from three types of observed foundation load test responses
(modified after Kulhawy 2004) B-21
Figure B16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footingsB-22
B-1
Figure B17 Measured load vs displacement curves for 3 shallow foundations on clay at Baytown Texas
(Stuedlein amp Holtz 2010)B-25
Figure B18 Characteristic stress vs square root of normalized displacement for all three foundations at
Baytown site (Stuedlein amp Holtz 2010) B-26
Figure B19 Normalized foundation stress vs square root of normalized displacements B-27
Figure B20 Applied foundation stress vs CPT-calculated stress for 67 footingsB-28
Figure B21 Applied footing stress vs CPT calculated stress on logarithmic scales B-29
Figure B22 Summary graph and equations of direct CPT method for evaluating the vertical B-30
Figure B23 Foundation soil formation parameter hs versus CPT material index Ic B-32
Figure B24 Bearing capacity ratio qmaxqnet versus CPT material index IcB-33
Figure B25 Influence factor for rectangular foundations from elastic theory solution (Mayne amp
Dasenbrock 2017) B-34
List of Tables
Table B1 Direct CPT methods for bearing capacity of footings on clean sandsB-4
Table B2 Direct CPT methods for bearing capacity of footings on clays B-9
Table B3 Summary of large footings soil conditions and reference sources of database B-17
Table B4 Sources for shallow foundation performace B-34
B-2
B1 DIRECT CPT METHOD FOR SHALLOW FOUNDATIONS
B11 Introduction
The analysis of shallow foundations on soils is classically handled as a two-step and two-part set of
calculations involving (a) bearing capacity commonly reliant on limit plasticity solutions and (b)
settlement or more appropriately termed displacements that are assessed via elastic continuum theory
The utilization of cone penetration tests (CPT) provides the necessary geotechnical data for the site-
specific information on the subsurface conditions at the project site including the geostratigraphy and
evaluation of input parameters This is still a viable approach where the CPT readings are interpreted to
give the soil unit weight (γt) effective friction angle (ϕ) undrained shear strength (su) preconsolidation
stress (σp) and elastic moduli (D and E) for analysis
An alternative approach is the use of CPT results to provide direct assessments of bearing capacity
andor settlements A comparison of these two distinctly different and alternate paths is depicted in
Figure B1
Figure B1 Conventional methods versus direct CPT approach to shallow foundation response
A number of available direct CPT approaches for shallow footings are reviewed within this report
Moreover as the entire load-displacement-capacity response of shallow foundations to loading occurs
as a continuous nonlinear phenomenon a single direct CPT solution is presented for the behavior of
B-3
footings on sands (drained) and clays (undrained) The methods discussed herein are specifically to
address the general case of vertical loading of shallow foundations Additional more complex situations
that consider load eccentricity moments inclined forces sloping ground and other facets may be
handled using well-established procedures that are discussed elsewhere (eg Vesić 1975 Kulhawy et
al 1983)
This report focuses on foundations on granular soils andor soils exhibiting drained behavior since
Federal Highway Administration (FHWA) studies have concluded that less than 1 of shallow
foundations for highway bridges are placed on clay soils (Paikowsky et al 2010) One reason for this is
because soft-firm intact clays require a more complicated process that assesses both a short-term
analysis (undrained) as well as long-term analysis (drained) thus requiring undrained bearing capacity
undrained distortion displacements drained bearing capacity and drained primary consolidation
settlements
B12 Direct CPT Methods for Bearing Capacity of Footings on Sands
Classical bearing capacity solutions are based most often in limit plasticity theory albeit can be
formulated from limit equilibrium cavity expansion and numerical methods Because the limit
plasticity solutions are prevalent and adopt total stress analysis they are usually developed as either
fully drained (Δu = 0) or fully undrained (ΔVV = 0) where Δu equals excess porewater pressure and
ΔVV equals volumetric strain Paikowski et al (2010) provides an extensive review of various solutions
There are a number of direct CPT methods for the evaluation of the bearing capacity of footings and
shallow foundations Table B1 lists a variety of these direct CPT approaches for footings on sands In
the direct approach a one-step process is used to scale the measured cone penetrometer readings (ie
measured cone tip resistance (qc) total cone tip resistance (qt) sleeve friction (fs) andor measured
porewater pressure u2)) to obtain the ultimate bearing stress (qult) of the foundation
Table B1 Direct CPT methods for bearing capacity of footings on clean sands
Method Surface Footing Remarks and Embedment Notes
Meyerhof (1956) qult = qc (B12) middotcw qult = qc (B12)(1+DeB)cw
Note for silty sand reduce by 05
cw = 10 dry or moist sand cw = 05 submerged sand
Meyerhof (1974) qult = qc (B + D)40 with stresses in tsf
Presumably dry sands B = fdn width (feet) and D = depth (feet)
Schmertmann (1978)
NA (applies to embedded footings)
Square qult =055σatm(qcσatm)078
Strip qult =036σatm(qcσatm)078
See Figure A2
Embedment applies De gt 05(1+B) for Blt1m De gt 12 m for B gt1m
B-4
Canadian qult = Rk0middotqc Applied to FS = 3 Geotech Society where Rk0 = 03 where FS = factor of (CFEM 199 2) safety
Tand et al (1995) NA qult = Rkmiddotqc + σvo See Figure A3 where Rk = fctn(De B)
Frank and Magnan (1995)
qult = Rk0middotqc where Rk0 = fctn(sand
consistency) Loose Rk0 = 014
Medium Rk0 = 011 Dense Rk0 = 008
qult = Rk1middotqc + σvo where Rk1 = function (De B L amp
sand consistency) Factor Rk1 = Rk0[1+Rk206+04(BL) (DeB)]
and Rk2 = 035 (loose) 050 (medium) and 085 (dense)
Loose sand qc lt 5 MPa
Medium sand 8 MPa lt qc lt 15MPa
Dense sand qc gt 20 MPa
Eslaamizaad amp Robertson (1996)
qult = KΦmiddotqc See Figure A4
See surface footing equation KΦ = function (BDe shape and density)
Lee amp Salgado (2005)
qbL = βbcmiddotqc(AVG)
where qc averaged over distance B
Not addressed See Figure A5 for factor βbc = fctn(B DR
K0 and sB)
beneath the base
Eslami amp Gholami (2005
2006)
See embedded footing solution
qult = Rk1middotqc where Rk1 = function (ratio DeB
and normalized qcσvo) See Figure A6
Measured qc and qcσvo are geometric means over 2B deep
beneath footing
Robertson amp Cabal (2007)
qult = KΦmiddotqc with KΦ = 016
See surface footing equation
Briaud (2007) qult = KΦmiddotqc with KΦ = 023
Based on full-scale tests at Texas AampM
Mayne amp Illingworth
(2010) qult = 018 qc-Mean
Note qc -Mean is averaged CPT cone resistance over depth of
influence z = 15 B deep
Based on 30 footing load tests on 12
sands
Lehane (2013) qult = 016 qc-AVE Note qc-AVE is averaged CPT cone resistance within z (m) = [B (m)
]07
Based on 47 load tests
Table B1 Continued
Notes B = minimum footing width (or diameter) L = footing length De = embedment depth cw = water
table correction σvo = total overburden stress at bearing elevation and σvo = the effective vertical stress
B-5
Figure B2 Direct bearing capacity relationship for embedded square and strip footings
situated on clean sands (after Schmertmann 1978)
Figure B3 Direct CPT bearing factor Rk as function of footing width B and embedment depth
from finite element analyses (Tand et al 1995)
B-6
Figure B4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio
and sand consistency (Eslaamizaad amp Robertson 1996)
B-7
Figure B5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp
Salgado (2005) in terms of footing size relative density and base settlement
Figure B6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and
normalized cone tip resistance (after Eslami amp Gholami 2005)
B13 Direct CPT Methods for Bearing Capacity of Footings on Clays
For direct CPT evaluations of foundation bearing capacity in clays Table B2 summarizes some of the
well-documented methods that are available Generally these assume that the loading is fast enough
and the permeability of the clay is also sufficiently low such that an undrained (ie constant volume)
condition is maintained This is fine for short term loading of foundations particularly for soft to firm
intact clays However the undrained case is not permanent Given sufficient time excess porewater
pressures in the clay will eventually dissipate and hydrostatic conditions will return to equilibrium
During this dissipation phase primary consolidation will occur that results in drained settlements Also
a drained bearing capacity condition will prevail As the CPT is advanced at a constant rate of 20 mms
the recorded readings normally constitute undrained behavior from a direct measurement viewpoint
Thus direct CPT methods are normally focused at an undrained foundation response Nevertheless the
drained footing case can be addressed using CPT results via use of conventional analysis approach and
effective stress limit plasticity solutions that require piezocone penetrometers with porewater pressure
measurements
B-8
B
D)
L
B4060(3501320kc
Many of the earlier solutions were based on data from mechanical penetrometers andor older electric
cones that measure the uncorrected tip resistance (qc) It is now well-recognized that this measurement
must be updated to the corrected cone resistance (qt) because of porewater pressure effects that act on
the mechanics of the load cell and geometry of the particular penetrometer design The results are
particularly significant in clays silts and mixed soils that are soft to firm to stiff where excess porewater
pressures are recorded
Table B2 Direct CPT methods for bearing capacity of footings on clays
Method Footing Comments NotesRemarks
Meyerhof (1974) qult = αbcmiddotqc
where 025 le αbc le 050
Applicable to saturated insensitive clays under short-term loading
Cone tip resistance
measured by
mechanical CPT
Trofimenkov (1974)
Approximation
qult asymp (qc33)09
(Assuming FS = 3)
Strip footings on clays and sandy clays 06m le B le 15m 1m le zemb le 25m
Mechanical CPT
with qc and qult in
kgcm2
Schmertmann (1978)
qult = function (qc and foundation shape)
Relationship shown in Figure A7 for square and strip footings
Cone tip resistance
measured by
mechanical CPT
Tand et al (1986)
qult = Rkmiddot(qc - σvo) + σvo
Factor Rk obtained from Figure A8 accounts for surface and deep footings Separate Rk factor for intact and jointed clays
= (qc1middotqc2)05 qc
where qc1 is geometric mean from bearing elevation to 05B deeper and qc2 is geometric mean from 05B to 15B beneath foundation base
Mix of data from
mechanical CPT qc
and electric CPT qc
LCPC Method
(Frank amp Magnan 1995)
Footing a surface
qult = kc middotqc + σvo
where kc = 032
Embedment case
Note Methods above generally consider undrained bearing capacity
B-9
Figure B7 Direct relationship between ultimate bearing stress in clays and measured cone tip
resistance (Schmertmann 1978)
B-10
Figure B8 Direct relationship between ultimate foundation bearing stress in clays and net
cone tip resistance (after Tand et al 1986)
B14 Direct CPT Assessments of Settlement and Displacements of Shallow Foundations
Methods for evaluating the magnitude of foundation displacements (s) can be found using
elastic theory subgrade reaction methods spring models and numerical simulations The
classical approach to footing settlement calculations on sands is to utilize elastic theory
solutions (eg Poulos amp Davis 1974) which take on the form
119902∙119861∙119868∙(1minus1205842) 119904 = B1
119864119904
where q = applied footing stress and I = displacement influence factor from elasticity theory
The value of I depends upon foundation geometry footing rigidity embedment depth variation of soil
modulus with depth compressible layer thickness and other factors as discussed by Mayne amp Poulos
(1999) For instance for a flexible circular footing of diameter B resting on an infinitely deep soil
formation having a constant modulus Es with depth the influence factor I is equal to 1 for the center
point The results of in-situ field tests such as pressuremeter tests (PMT) standard penetration tests
(SPT) cone penetration tests (CPT) flat dilatometer tests (DMT) and other methods can be used to
ascertain the input value of soil modulus Es
Alternatively direct CPT methods have been investigated to provide a one-step assessment of
foundation displacements Selected methods for direct CPT evaluation of shallow foundation
settlements on granular soils is include DeBeers amp Martens (1957) Meyerhof (1965) DeBeer (1965)
Thomas (1968) Schmertmann (1970) Meyerhof (1974) Berardi amp Jamiolkowski amp Lancellotta (1991)
Robertson (1991) Lutenegger amp DeGroot (1995) Lehane amp Dougherty amp Schneider (2008) Gavin amp
Adekunte amp OrsquoKelly (2009) Mayne amp Illingworth (2010) and Uzielli amp Mayne (2012)
B141 Large Scale Footing Load Tests
The measured load-displacement response of shallow foundations is conducted in the field using either
stepped loading procedures or continuous rate of displacement methods Results from the well-
documented test program at the Texas AampM University (TAMU) site (Briaud amp Gibbens 1999 Briaud
2007) for five spread footings on sand are shown in Figure B9 Each footing clearly shows a nonlinear
B-11
behavior to loading over the range of testing This site is one of the national geotechnical
experimentation sites (NGES) funded by the National Science Foundation (NSF) FHWA and American
Society of Civil Engineering (ASCE)
Figure B9 Measured load-displacements for each of the five large footings at TAMU (Briaud
amp Gibbens 1994) sand site
B142 Generalized Direct CPT Method for Footing Response on Soils
The use of applied foundation stress versus normalized displacement curves (q vs sB) is generalized to
footings on sands silts intact clays and fissured clays A square root plotting of the normalized
displacements (sB) permits a single parameter characterization for each specific soil type that in turn
correlates with the net cone tip resistance The generalization is based on a statistical review of data
from large foundations (B gt 05m) involving 70 full-scale load tests with available CPT data For fine-
grained soils the data are from electronic piezocone tests where the proper correction from qc to qt has
been obtained to allow the best possible results The methodology permits an evaluation of the load-
displacement-capacity response of shallow square footings based on CPTs performed in the four types
B-12
of ground conditions For the TAMU footings the characteristic stress-normalized displacement
response is shown in Figure B10 where all five foundations can be represented by a single curve
Figure B10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU
(Briaud amp Gibbens 1994) footings
B143 Characteristic Stress-Displacement Curves
Fellenius amp Altaee (1994) recommended the concept of a characteristic stress (q) vs normalized
displacement (sB) curve for foundation response on a given soil formation later supported by Decourt
(1999) Briaud amp Gibbens (1999) Lutenegger amp Adams (2003) and Briaud (2007) In this approach the
measured load (Q) vs displacement from individual footings of various sizes that rest on the same soil
conditions collapse to a unified relationship given in Equation B2
119939119943 119956 119954119938119953119953119949119946119942119941 = 119938119943 (119913
) B2
B-13
where af and bf are empirical fitting coefficients (Decourt 1999 Uzielli amp Mayne 2011 2012)
In a review of measured load tests data from large spread footings situated on different sands it has
been suggested that Equation B2 can be reduced to a single parameter expression by use of square root
plotting where the exponent bf is equal to 05 (Mayne amp Illingworth 2010 Mayne et al 2012)
119956 119954119938119953119953119949119946119942119941 = 119955119956radic
119913 B3
where rs = fitted soil parameter from best fit line regression In the case of sands the coefficient rs
represents the supporting soil conditions including particle size relative density and fines content as
well as geologic origin aging and overconsolidation effects
The results from the five TAMU footings are presented in this format in Figure B11 that shows the
formation factor rs = 486 MPa for this sand deposit This single coefficient can be used to express the
nonlinear load versus settlement of any size footing on this sand site Moreover one criterion from
Eurocode for bearing capacity that is used by the European community identifies that stress (or load)
which corresponds to (sB) = 10 That is when the displacement equals ten percent of the footing
width then the capacity has been reached Therefore adopting this criterion for footings on sands
the ultimate bearing stress (qult) for footings on sand can be taken when (sB) equals 010 or when
square root of (sB) equals 0316 In that case this gives qult equal to 0316 middot rs For the TAMU site this
gives qult equal to 0316 times (486) which equals 153 MPa as illustrated by Figure B12
B-14
Figure B11 Characteristic stress versus square root normalized displacements for TAMU
footings
B-15
Figure B12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sand
B15 Footing Database
A database of spread footings and large plates was assembled which considers only full-size shallow
foundations (05 le B le 6 m) that rest on sands silts and clays Sites for the foundations were also subjected to CPT The inclusion solely of large footings are significant because results from small-scale
model footings exhibit scale effects In fact experimental centrifuge work and numerical simulation
studies have shown that bearing capacity factors are size-dependent especially for factor Nγ decreasing
with footing width B (Kimura et al 1985 Cerato amp Lutenegger 2007 Mase amp Hashiguchi 2009) A
direct approach based on full-size footings provides a more reliable means for foundation evaluation
The compiled database is given in Table B3 with a total of 70 large footings and plates These include a
listing of 34 foundations on 13 sands 11 footings on 4 silt deposits 13 footings or large plate load tests
on 6 intact clays and 12 foundations on 5 fissured clays Most of the footings were square or nearly
square (80) while the remaining were circular The largest foundation was a mat 5 m by 14 m in plan
loaded by stacking Kentledge concrete blocks to cause a bearing failure in the underlying soft clay For
sands the largest footing consisted of the 6-m square concrete pad In terms of equivalent square
footings mean footing sizes were 155 m (sands) 114 m (silts) and 126 m (clays)
B-16
Additional details on the individual load tests and site conditions are given elsewhere (Mayne 2009
Mayne amp Illingworth 2010 Uzielli amp Mayne 2011 2012 Mayne et al 2012 Mayne amp Woeller 2014)
Table B3 Summary of large footings soil conditions and reference sources of database
Sand Site Location Soil Conditions Footings Numbers
Shapes and Sizes
ReferencesSource
College Station
Texas Pleistocene sand 5 Square 1 15 25 33 m Briaud amp Gibbens (1999)
Kolbyttemon Sweden Glaciofluvial sand 4 Rect B = 06 12 17 24 m
Bergdahl et al (1985 1986)
Fittija Sweden Glaciofluvial sand 3 Rect B = 06 17 24 m Bergdahl et al (1984 1985)
Alvin West Texas Alluvial sand 2 Circular D = 235 m Tand et al (1994)
Alvin East Texas Alluvial sand 2 Circular D = 22 m Tand et al (1994)
Perth Australia Silceous dune sand 4 Square B = 05 and 10 m
Lehane (2008)
Grabo T1C Sweden Compacted sand fill
1 Square B = 046 m Long (1993)
Grabo T2C Sweden Compacted sand fill
1 Square B = 063 m Long (1993)
Grabo T3C Sweden Compacted sand fill
1 Square B = 080 m Long (1993)
Labenne France Dune sand 6 Square B = 07 and 10 m
Amar et al (1998)
Green Cove Florida Brown silty sand 1 Circular D = 182 m Anderson et al (2006)
Durbin South Africa
White fine sand 1 Square B = 609 m Kantley (1965)
Porto Portugal Residual clayey sands
1 Circular D = 12 m and 1 plate
Viana da Fonseca (2003)
Jossigny France Soft clayey silt 2 Square B = 1 m Amar et al (1998)
Tornhill Sweden Glacial Baltic till 3 Square B = 05 1 2 m Larsson (2001)
Vagverkel Sweden Stiff medium silt 3 Square B = 05 1 2 m Larsson (1997)
Vattahammar Sweden Brn-Gry layered silt 3 Square B = 05 1 2 m Larsson (1997)
B-17
Table B3 Continued
Clay Site Location Soil Conditions Footings Numbers Shapes and Sizes
ReferencesSource
Baytown Texas Fissured Beaumont 2 plates with d = 076 m Stuedlein amp Holtz (2008)
Belfast Ireland Soft clay silt ldquosleechrdquo
1 Square Pad B = 20 m Lehane (Geot Engr 2003)
Bothkennar Scotland Soft silty clay 2 Square Pads B = 22 24 m
Jardine et al (1995)
Bangkok Thailand Soft to stiff clay 4 Square 067 075 090 10 m
Brand et al (1972)
Haga Norway Stiff OC clay 2 Square Footings B = 10 m
Andersen amp Stenhammar (1982)
Rio Grande Brazil Sandy residual clay 3 Square 04 07 and 10 m
Consoli et al (1998)
Shellhaven England Soft estuarine clay Rectangular 5 m by 14 m Schnaid et al (1993)
Texas City A Texas Coast Fissured Beaumont 3 circular plates d = 058 m
Tand et al (1986)
Texas City B1 Texas Coast Fissured Beaumont 2 circular plates d = 058 m
Tand et al (1986)
Texas City B2 Texas Coast Fissured Beaumont 1 circular plate d = 058 m
Tand et al (1986)
Alvin Texas Texas Coast Fissured Beaumont 3 circular plates d = 058 m
Tand et al (1986)
For the series of footings on sands Figure B13 shows the summary of measure formation factors (rs) for
the 34 foundation load tests Also indicated on the graph are the corresponding mean qc values of the
sands averaged over 15middotB beneath the bearing elevations
Note also in sands that the qt and net resistance (qnet) are all very close to the measured qc since (a)
porewater pressure corrections for the a-net value are negligible in clean sands and (b) overburden
stresses are typically only 1 to 3 of the magnitude of qc This is especially true of shallow depths
Therefore for clean sands qnet is approximately qt which is approximately qc
B-18
Figure B13 Summary of sand formation factors with corresponding CPT resistances
As suggested by Briaud (2007) various in-situ measurements can be used to normalize the results from
cone penetration tests in this case Herein the qc from the CPT soundings in the sands are used to
normalize the footing stress axis as presented in Figure B14 It is evident that the results of the load-
displacement response of all 34 footings on 13 different sands can be captured by the characteristic
stress versus normalized displacement with the inclusion of the site-specific cone tip resistance In this
case the mean trend for all footings and sand sites indicates a simple expression shown in Equation B4
119954119943119952119952119957119946119951119944 119956 Clean quartz-silica sands = 120782 120787120790120787radic B4
119954119940 119913
B-19
Figure B14 Normalized foundation stresses vs square root of normalized displacement for
spread footings on sand (modified after Mayne et al 2012)
The results of measured force-displacement curves (Q vs s) from footing load tests are nonlinear and
thus the definition of capacity must be addressed Kulhawy (2004) identified three main curve types
that occur during load testing depicted in Figure B15 The most common response (Type A) shows no
clear peak and is observed in sands and silts as well as slow loading of insensitive clays and fissured
clays In this case ldquocapacityrdquo can be defined by the Eurocode corresponding to the load (or stress) when
sB=10 (Amar et al 1998) For foundations situated on many clays a plateau capacity is reached as
depicted as Type B response In rare cases where sensitive clays or structured soils are encountered a
Type C relationship occurs whereby the load-displacement curve reaches a peak followed by strain
softening In the one case observed in this study for Type C loading (Haga clay) the corresponding
ldquopeak capacityrdquo was reached at a pseudo-strain of sB = 4 as shown in Figure B16 followed by some
strain softening In the cases of soft clays at Belfast and Bothkennar a plunging type (plateau failure)
was achieved at around sB=7
B-20
Figure B15 Definitions of capacity from three types of observed foundation load test
responses (modified after Kulhawy 2004)
B-21
Figure B16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footings
In the case of fine-grained soils that develop excess porewater pressure during cone penetration the
use of the net total cone resistance (qtnet) which equals the cone tip resistance minus the total stress
must be considered (Lunne et al 1997) As noted earlier the correction of CPT data in sands is minimal
because of the low value of penetration porewater pressures and the fact that total overburden stress is
small relative to the cone resistance especially for shallow soundings Thus the use of qtnet may be
substituted for qc into the trend of Figure B1
B16 Undrained and Drained Capacity
The footings on sands typically show drained response with no excess porewater pressures developed
during loading For the foundations situated on silts analyses by Larsson (1997 2001) concluded that
these too show drained conditions For shallow foundations on clays the entire range of drainage cases
are possible ranging from undrained to partially-drained to fully-drained depending upon the applied
rate of loading relative to the permeability of the geomaterial If sufficient data were available for each
of the clay sites then the recent evaluation of drainage condition could be assessed using normalized
velocity criteria (Chung et al 2006) However this was not possible with the current data set Instead
B-22
the test loadings of footings on clays have essentially been assumed to primarily occur under undrained
conditions following recommendations by Tand et al (1986) For this case a limiting bearing stress can
be calculated from bearing capacity analyses and related directly to the CPT readings
From classical bearing capacity calculations using limit plasticity solutions the ultimate foundation stress
is shown in Equation B5
= 119925119940 ∙ 119956119958 B5 119954119958119949119957
where Nc equals the bearing factor for constant volume (514 for strip and 614 for square or circular
plan) and su equals the undrained shear strength of the clay For the case of soft to firm clays of low to
medium sensitivity the operational strength can be related to the preconsolidation stress (Mesri 1975
Jamiolkowski et al 1985) as shown in Equation B6
prime 119956119958 = 120782 120784120784120648119953 B6
Finally the effective preconsolidation stress has been linked directly to the net cone resistance (Chen amp
Mayne 1996 Demers amp Leroueil 2002) as shown in Equation B7
prime 120648119953 = 120782 120785120785(119954119957 minus 120648119959119952) B7
Combining Equations (B5) (B6) and (B7) results in direct expressions for undrained bearing capacity on
intact clays from CPT results as shown in Equations B8 and B8-2
Strip footing 119954119958119949119957 = 120782 120785120789120785(119954119957 minus 120648119959119952) B8
B-23
Squarecircle 119902119906119897119905 = 0445(119902119905 minus 120590119907119900) B8-2
Equation (B8-2) is in excellent agreement with Tand et al (1986) database study for the case of intact
clays and footings with no embedment Their study also provided a lower relationship for foundations
situated on jointed clays specifically recommending that the bearing stress not exceed 030 qtnet
Review of the available data in Figure 3 shows this to be rather conservative and a more realistic value
may be taken at 040 qtnet for fissured clays
B161 Normalized Undrained Footing Response
The characteristic stress vs normalized displacement curves for footings on clays exhibiting undrained
conditions can be verified using a relatively recent case study on Beaumont clay by Stuedlein amp Holtz
(2010) The Baytown Texas site was investigated using an exploration program of soil borings
piezocone soundings and laboratory testing The field testing included a large square footing (B = 276
m) and two small circular plates (D = 076 m) The measured load-displacement responses for all three
foundations are presented in Figure B17 Of course the large footing carried considerably higher loads
than the two small plates on the same soil deposit Yet using the aforementioned characteristic stress
(q) vs square root of normalized displacement (sB) essentially gave a unique relationship for all 3
foundations as presented in Figure B18 In this case utilizing the form of Equation B2 the
characteristic coefficient rs is 177 MPa for this deltaic clay formation
B-24
Figure B17 Measured load vs displacement curves for 3 shallow foundations on clay at
Baytown Texas (Stuedlein amp Holtz 2010)
B-25
Figure B18 Characteristic stress vs square root of normalized displacement for all three
foundations at Baytown site (Stuedlein amp Holtz 2010)
B17 General Direct CPT Method
In a manner analogous to the normalization of applied foundation stresses shown in Figure B14 for
footings on sands a generalized direct CPT method for shallow foundations on different soils can be
made in Equation B9
120782120787 119956 119954 = 119945119956 ∙ 119954119957119951119942119957 ∙ (119913
) lt 119954119940119938119953119938119940119946119957119962 B9
where qcapacity is equal to the foundation bearing capacity of the ground and hs is equal to the empirical
fitting term that depends on soil type Specifically hs equals 058 for sands 112 for silts 147 for
fissured fine-grained soils and 270 for clays A summary of the normalized stress versus the normalized
displacement curves for these four soil types is presented in Figure B19 where the data fitting to obtain
the hs parameter on clays are limited to (sB) lt 004 corresponding to undrained loading The data on
B-26
silts and sands are considered fully drained whereas the fissured clay subset may be partially drained to
undrained The statistical measures for obtaining the fitted parameter hs are quite good as shown in
Figure B20 with the coefficient of determinations (r2) values of 0947 for sands 088 for silts 0935 for
fissured and 0925 for intact clays Additional statistical evaluations on the database are provided by
Uzielli amp Mayne (2011)
Figure B19 Normalized foundation stress vs square root of normalized displacements
B-27
Figure B20 Applied foundation stress vs CPT-calculated stress for 67 footings
The same data are shown on Figure B21 in a log-log plot to emphasize the early parts of the load-
displacement responses of these footings The best fit line from regression analyses shows excellent
statistical correlations As detailed earlier the undrained bearing capacity of square or circular footings
on clays can be taken as approximately 040 times qtnet For silts and sands that experience drained
loading with no excess porewater pressures being developed and no clear peak value for capacity the
(sB) =10 criterion can be used In this case Equation 7 can be used directly to obtain stress q equal to
qcapacity when (sB)05 is 0316
An alternate approach is presented in Figure B22 summarizing the direct CPT method for estimating the
load-displacement-capacity of shallow square footings on four soil types The maximum values of soil
bearing capacity (qmax) can be capped at stresses corresponding to a percentage of the average
measured qtnet determined over the depth interval from the foundation bearing elevation to 15middotB
below the foundation In such an approach the foundation bearing capacity can be taken as
(qcapacityqtnet) of 02 for sands 035 for silts 040 for fissured and 045 for intact clays Using the assigned
values of the hs parameter the corresponding cut-off for the general stress-normalized displacement
relationship given by Equation 7 can be established specifically giving corresponding values of the
B-28
normalized displacements which can also be considered as pseudo-strains equal to (sB) max of 4 for
clays 7 for fissured clays 10 for silts and 12 for sands
Figure B21 Applied footing stress vs CPT calculated stress on logarithmic scales
B171 CPT Soil Material Index Ic
The soil behavioral type can be assessed indirectly by CPT using a material index (Ic) as defined by
Robertson (2009a b) shown in Equation B10
119920119940 = radic[120785 120786120789 minus 119949119952119944119928119957119951]120784 + [120783 120784120784 + 119949119952119944119917119955]120784 B10
B-29
Figure B22 Summary graph and equations of direct CPT method for evaluating the vertical
where Qtn equals the stress-normalized cone tip resistance and Fr equals the normalized sleeve friction
calculated from the cone penetrometer readings as shown in Equations B11 and B12 respectably
(119954119957minus120648119959119952)120648119938119957119950 = 119951 B11 119928119957119951 prime (120648119959119952120648119938119957119950)
119943119956 119917119955() = 120783120782120782 ∙ B12 (119954119957minus120648119959119952)
where σatm equals a reference stress equal to atmospheric pressure (σatm = 1 atm asymp 1 bar asymp 100 kPa) In
the initial evaluation the exponent n is set to n = 1 to find the soil behavioral type (SBT) based on a 9-
zonal chart as shown in Figure 23 The value of exponent n varies with soil type ranging from n = 1 in
intact clays and decreasing with increasing grain size to around approximately 075 in silts and
B-30
approximately 05 plusmn 02 in clean sands The appropriate value of exponent n is found by iteration until
conversion using the relationship (Robertson 2009b) shown in Equation B13
prime 120648119959119952 119951 = 120782 120785120790120783 ∙ 119920119940 + 120782 120782120787 ( ) minus 120782 120783120787 le 120783 120782 B13 120648119938119957119950
As the distinction that footings on intact clays subjected to fast loading are behaving primarily under
undrained conditions (ie constant volume) Robertson (2009a) suggested that this occurs when Ic gt 27
while in contrast Ic lt 25 corresponds more or less to drained loading (ie no excess porewater
pressures)
The CPT material index can be used to identify soil type and the corresponding characteristic hs value for
estimating footing load-displacement response In a number of the case studies investigated the results
of the sleeve friction readings were also available to allow the calculation of Ic with depth at these sites
Figure B23 shows the tentative relationship between the values of the hs parameter plotted versus the
corresponding CPT material index with an approximate trend given by Equation B14
120784120785 119945119956 = 120784 120790 minus 120783120787 B14
119920119940 120783+( ) 120784120786
B-31
Figure B23 Foundation soil formation parameter hs versus CPT material index Ic
Soil behavioral type from the Ic ranges given by Robertson (2009b) and are shown in Figure B23 with an
overall general agreement with the known soil classifications In summary the CPT can provide an
evaluation of the load-displacement-capacity response directly via Equation B15 which may be of
interest in mixed soil types such as silty sands sandy silts and the like
120784120785 Footing stress 119954 = 119954119951119942119957 ∙ radic(119956frasl119913) ∙ [120784 120790 minus
)120783120787] B15
120783+(119920119940frasl120784120786
As discussed earlier foundation bearing capacity may be taken by a limiting value of pseudo-strain (sB)
max or by qmax as specified as a percentage of qtnet This definition of bearing capacity can be seen in
comparison to soil type as seen in Figure B24
B-32
Figure B24 Bearing capacity ratio qmaxqnet versus CPT material index Ic
B172 Rectangular Foundations
This same elastic solution from Giroud (1968) for a CPT method for square and circular foundations were
utilized for rectangular foundations The study covered rectangular distortions (AB) ranging from 1
(square) to very long foundations with AB equal to 20 where A represents the foundation length and B
represents the foundation width (Mayne amp Dasenbrock 2017) The influence factor for rectangular
shaped foundations is given in Figure B25 and the expression in Equation B16
= (119912119913)120782120785120786120787 119920119912119913 B16
B-33
Figure B25 Influence factor for rectangular foundations from elastic theory solution (Mayne
amp Dasenbrock 2017)
Including 32 full scale load tests on sands results from an additional 98 shallow foundations have been
reported in previous studies The foundations were primarily square or circular and among these include
large spread footings with measured settlements and bearing capacity Table B4 shows a summary of
the number of footings and footing sizes used in the study The range of length to width ratios varied
from 1 lt (AB) lt 23 with an average value of (AB) equal to 238 The embedment depth to footing
width ranged from 0 lt DeB lt 222 and averaged 046
Table B4 Sources for shallow foundation performace
Source of Data No of
Footings
Mean B
(m)
Max B
(m)
Min B
(m)
Mean A
(m)
Max A
(m)
Min A
(m)
Mayne et al (2012) 32 149 609 046 149 609 046
Lehane (2011) 8 041 027 060 041 060 027
Gifford et al (1987) 17 846 1588 527 1591 3596 701
Jeyapalan amp Boehm
(1986)
13 1022 2892 400 1671 3049 640
B-34
Schmertmann
(1970)
31 945 5608 061 1288 8668 061
Papadopoulos
(1992)
29 870 3600 100 1300 7290 100
Total 130 671 5608 027 1012 8668 027
The solution for shallow foundation settlements given back in Equation B1 combined with
the expression in Equation B16 takes on the form
119954∙119913∙119920119912119913∙(120783minus120642120784) 119956 = B17
119916119956
Combining the direct CPT equation developed from the 32 full scale load tests (Equation B15)
together with the influence factor for rectangular shaped foundations becomes
minus120782120785120786120787 120784120785 119912 Footing stress 119954 = 119954119951119942119957 ∙ radic(119956frasl119913) ∙ [120784 120790 minus
)120783120787] ∙ [ ] B18
120783+(119920119940frasl120784120786 119913
B18 Conclusions
A direct CPT method for square rectangular and circular shallow footings is developed using a database
of 166 full-scale field load tests where the minimum footing size of an equivalent square width of 05 m
is established to avoid scaling issues The use of load vs displacement is generalized by characteristic
stress vs normalized displacement (sB) and further simplified by a square root plotting technique that
captures the ground response in a single parameter that is related to the qtnet Data are grouped
according to four main soil categories sands silts fissured clays and intact clays It is generally believed
that the footings on sands and silts exhibit fully drained behavior while in intact clays an undrained
response occurs under conditions of constant volume
B-35
The summary equation for evaluating the vertical stress-displacement-capacity of square rectangular
and circular footings is given by
minus120782120785120786120787 119956 119912 119954119943119952119952119957119946119951119944 = 119945119956 ∙ 119954119957119951119942119957 ∙ radic ∙ ( ) lt 119954119950119938119961 B19
119913 119913
where the parameter hs also relates directly to Ic The foundation capacity depends upon the mode of
failure as well as drainage conditions and can be taken as a fraction of qtnet where qmaxqtnet is 020 in
sands 035 in silts 040 in fissured clays and 045 in intact clays as shown in Figure 24 Alternatively the
capacity can be defined using a limiting pseudo-strain given by (sB) max of 4 in clays 7 in fissured
10 in silts and 12 in sands
B-36
APPENDIX C
List of Figures
Figure C1 Components of Axial Pile Capacity C-3
Figure C2 Selected alpha relationships as function of the undrained shear strengthC-5
Figure C3 Pile side friction coefficient β in terms of effective friction angle and overconsolidation ratio
C-11
Figure C4 Concept of Direct versus Rational Method for CPT evaluation of axial pile capacityC-14
Figure C5 Soil behavior type using CPT via original UniConeC-19
Figure C6 Soil behavior type using CPT via Modified UniCone C-19
Figure C7 Nine-zone soil behavioral soil type using normalized piezocone parameters C-21
Figure C8 Summary of Modified UniCone Method for direct CPT assessment of axial pile capacity C-22
Figure C9 Elastic continuum solution for axial pile response under compression and tension loadingC-24
List of Tables
Table C1 Alpha Methods for Axial Pile Side Friction where fp = αmiddotsu C-5
Table C2 Beta Methods for Axial Pile Side Friction where fp = βmiddotσvoC-9
Table C3 Selection of Direct CPT Methods for Axial Pile CapacityC-15
C-1
C1 EVALUATING DEEP FOUNDATION RESPONSE FROM CONE
PENETRATION TESTS
C11 Introduction
The axial response of deep foundations includes the load-displacement-capacity and axial load transfer
when driven pilings and drilled shafts are loaded in compression and uplift The use of cone penetration
testing (CPT) especially seismic piezocone tests (SCPTu) are advantageous since they provide
information on the subsurface soils and their geomaterial properties The SCPTu provides at least four
measurements on soil behavior with depth including (a) cone tip resistance qt (b) sleeve friction fs (c)
penetration porewater pressure u2 and (d) shear wave velocity Vs This offers the opportunity to
determine the geostratigraphy unit weight effective overburden stress shear strength stress state
and stiffness of the ground from a single exploratory sounding thus values in economic expediency
and reliability The axial pile capacity can be calculated based on static equilibrium of forces acting along
the sides and base of the pile foundation The displacement of the pile can be ascertained using
elasticity theory from closed-form solutions boundary elements andor finite element analyses Load
transfer occurs along the length of the pile and only a portion of axial forces are transmitted to the base
or toe or tip of the pile These too can be assessed using elasticity solutions
An alternate approach to pile capacity involves the utilization of direct CPT methods available
since the 1970s but in the past 10+ years a number of new and statistically reliable algorithms
have been developed which can be implemented for highway design and construction
C111 Axial Pile Capacity
The axial compression capacity of a single pile foundation is composed of a shaft or side component and
end-bearing component at the base as depicted in Figure 1 For a circular pile the side capacity (Qs) is
determined from the unit side friction (fp) acting along the surface area of the shaft which is As = πmiddotdmiddotL
where d = pile diameter and L = length embedded below grade
If the magnitude of side friction is uniform and constant with depth the side capacity is simply
C-2
119928119956 = 119943119953 middot 119912119956 C1
Moreover however many piles extend through multiple layers and a summation of unit side
frictions acting on various pile segments must be tabulated over the length of the pile as
suggested by Figure C1
Figure C1 Components of Axial Pile Capacity
The unit-end bearing resistance (qb) acts over the base of the pile tip where the area of a circular pile is
given by Ab = πmiddotd24 For piles in compression loading the base capacity is determined from Equation
C2
119876119887 = 119902119887 middot 119860119887 C2
and for piles in tension (or uplift) it is normally taken that Qb = 0
C-3
C112 Pile Unit Side Friction
Several different approaches can be adopted for evaluating the unit pile side friction (fp) prior to full-
scale load testing and construction (Poulos amp Davis 1980 ONeill 2001) The most common types
including the alpha and beta methods as summarized in Tables 1 and 2 respectively In the alpha
method an empirical coefficient (α) is applied to the undrained shear strength (su) of clay soils to
obtain
119891119901 = 120572 middot 119904119906 C3
An illustrative example of alpha curves is shown in Figure C2 A difficulty with the alpha
method is the evolution of many variants and changes to the expressions for its estimation It is
also restricted in its use for specific types of piles in clays and fine-grained soils
C-4
Figure C2 Selected alpha relationships as function of the undrained shear strength
Table C1 Alpha Methods for Axial Pile Side Friction where fp = αmiddotsu
C-5
Reference
Source
Alpha Equation Remarks
Tomlinson
(1957)
2716801102
uu ssAll pile types (steel
concrete timber)
Note su in ksf
Tomlinson
(1957)
2616201102
uu ssConcrete piles
Note su in ksf
McClelland
(1974)
Kerisel α = 07 - 031middotln(su)
Peck α = 10 - 02middotsu
Woodward α = 091middot(su)091
Driven piles in clay
Note su in ksf
Semple
(1980) )1(511
)1(1
OCR
OCR Summary from 9 series of
pile load tests
American
Petroleum
Institute (API
1981)
For suσvo le 035 α = 10
For suσvo gt 080 α = 05
Otherwise α = 1 + 1111middot (035 - suσvo)
Driven steel pipe piles
Tomlinson
(1986)
1 α = 55 (su)-091
2 α = 785 (su)-102
3 α = 605 (su)-100
where 75 lt su (kPa) lt 210
1 Driven piles
2 Bored piles
3 Driven piles in till with L
lt 10middotd
API (1987) 1 α = 10 for su lt 25 kPa Driven piles in clay other
than Gulf of Mexico
C-6
2 α = 125 - 001middotsu for 25 lt su lt 75 kPa
3 α = 050 for su gt 75 kPa
Kulhawy amp
Jackson
(1989)
α = 021+026middot(σatm su) le 1 Analyses of 106 drilled
shaft foundations
Table C1 Continued
API (1989) 05 For suσ vo le 1 α =
su radic frasl prime σvo
minus025 su For suσ vo gt 1 α = 05 ∙ ( frasl prime ) σvo
Driven steel pipe piles
Kulhawy amp
Jackson
(1989)
OCR
OCR
sin50
tan)sin1( sin
Λ = (1-CsCc) asymp 080 for
insensitive clays
Nowacki et al
(1996)
minus05 su For suσvo le 07 α = 05 ∙ ( frasl prime ) σvo
minus02 su For suσvo gt 07 α = 055 ∙ ( frasl prime ) σvo
Driven pile foundations
Kolk and van
der Velde
(1996)
α =09middot FL middot (suσvo)-03 lt 10
FL = [ (L - z)d]-02 = length term
L = pile length
Note su obtained from
UU lab tests
Applicable to driven piles
in clays
C-7
Knappett amp α = 1 for su le 30 kPa Non-displacement piles in
Craig (2012) fine-grained soils α = 116 - (su185) for 30 kPa le su le 150 kPa
α = 035 for su gt 150 kPa
Knappett amp
Craig (2012) α = 055 ∙ 02 40
( ) (LfraslD)
∙ su ( frasl prime σvo
minus03
) Displacement piles in fine-
grained soils
z = depth at point considered
d = outside diameter of pile
Karlsrud et al
(2005 NGI
Method)
α = function (suσvo and plasticity index) Driven pilings
Fleming et al
(2009) 05 minus05 su su For su σvo le 1 120572 = ( frasl prime ) ∙ ( frasl prime )
σvo NC σvo
05 minus025 su su For su σvo gt 1 α = ( frasl prime ) ∙ ( frasl prime ) σvo NC σvo
For driven piles in clay
where (suσvo)NC is the
normalized shear strength
for normally-consolidated
clay
Brown et al
(2010)
α = 0 for 0 lt z le 5 feet
α = 055 for z gt 5 feet and (suσatm) le 15
α = 055-01middot(suσatm -15) for 15le (suσatm) le 25
Drilled Shafts and Bored
Piles
Table C1 Continued
C-8
OCRK )sin1(0
Karlsrud
(2012)
α = function (su σvo and plasticity index) Driven pilings
Note as reported by Karlsrud (2012)
In the beta method the coefficient β is applied to the effective overburden stress (σvo) at the point of
concern along the pile length
prime = 120573 C4 119891119901 middot 120590119907119900
Table C2 Beta Methods for Axial Pile Side Friction where fp = βmiddotσvo
Reference Source Beta Equation Remarks
Burland (1973) β = (1-sinϕ) middot tan ϕ Piles in NC clays
Meyerhof (1976) β = Kmiddottanδ where K = K0 = for NC clays
K = 15middotK0 for OC clays
Poulos amp Davis
(1980)
β = (1-sinϕ) middot tan ϕ middot OCR05 Piles in OC stiff clays
Table C2 Continued
C-9
Kulhawy amp Jackson
(1989)
β = (1-sinϕ) middot tan ϕ middot OCRsinϕ Piles in quartz sands and insensitive
clays
ONeill (2001) β = (1-sinϕ) middot tan θ middot OCRsinϕ Drilled shafts where θ = (δϕ)middotϕ and
δ = interface friction between pile and soil
Fleming et al
(2009)
β = Kmiddottanδ K = lateral stress coefficient and δ =
friction angle between soil and pile
material
Karlsrud (2012) β = fctn (OCR and PI) PI = plasticity index of the clay ()
Mayne and Niazi
(2017)
β = CM middot CK middot K0 middot tanϕ
where K0 = (1-sinϕ) middot OCRsinϕ for soils
that are virgin loaded then unloaded
CM = pile material factor = 1 (drilled
augered) 09 (prestressed or precast
concrete) 08 (timber) and 07
(steel) and CK = pile installation factor
= 09 (bored or augered) 10 (low
displacement eg H-pile or open-end
pipe) and 11 (driven high
displacement eg prestressed
concrete closed-end pipe)
As the beta approach applies to all types of soils (gravels sands silts and clays) and more or
less has remained unchanged since its advent circa 1970 it has been selected for further
discussion herein In the direct form for calculation of unit pile side friction the expression is
given by
119874119862119877119904119894119899120601 prime prime 119891119901 = 119862119872 ∙ 119862119870 ∙ (1 minus 119904119894119899120601prime) middot ∙ 119905119886119899120601prime ∙ 120590119907119900 C5
C-10
where CM is a soil-pile interface friction coefficient and CK = pile installation factor Values of CM are in
the range 07 lt CM lt 10 depending upon pile material (steel wood concrete) and ranges for CK vary
09 lt CK lt 11 depending upon method of installation (auger drill driven) as detailed in Table 2
The value of K0 is limited to the passive stress coefficient which for the simple Rankine case is given by
KP = (1+sin ϕ) (1 - sin ϕ ) Thus there is a maximum value of overconsolidation ratio for which
Equation C5 applies given by OCRlimit = [(1+sin ϕ ) (1 - sin ϕ )2] (1sinϕ)
The corresponding graph in Figure C3 shows the relationship for β in terms of ϕ and OCR
Figure C3 Pile side friction coefficient β in terms of effective friction angle and overconsolidation ratio
C113 Pile Unit End Bearing
C1131 Theoretical Considerations
The evaluation of the end bearing resistance of pile foundations is commonly determined using
limit plasticity theory where
prime Drained loading = C6 119902119887 119873119902 middot 120590119907119900
C-11
Undrained loading 119902119887 = 119873119888 middot 119904119906 C7
where the value of σvo is calculated at the depth z = L and the value of su is the average undrained shear
strength from z = L to a depth z = L + d beneath the pile tip For a circular pile the limit plasticity solution
for undrained loading gives (Vesic 1977) a value Nc = 933 while a deep strip foundation would employ a
value of Nc = 824 For drained loading the expression for Nq for a deep foundation is given by (Vesic
1977)
)]arctan()sin1(tan21[)](tan1[sin1
sin1)tanexp( 2 BLABNq
C8
where A and B are the pile plan dimensions (for a circular pile A = B) and L = pile length For a
circular pile this can be approximated by
asymp 077(120601prime75deg) 119873119902 C9
over a range of effective friction angles 20deg le ϕ le 45deg
C1132 Practical Considerations
For drained loading of sands the full calculated end bearing capacity will never be realized because it
would require the pile to move a distance equal to its diameter This is beyond practical use and
therefore the limit plasticity solutions must be clipped to a fraction of the calculated value Another
reason for using a reduced end-bearing capacity is due to strain incompatibility since the side capacity is
mobilized early while the end-bearing is engaged much later So to have compatible values of Qs and
Qb the qb must be reduced using the following guidelines (Randolph 2003 Mayne 2007)
prime prime C10 119902119887 = 119873119902 ∙ 120590119907119900 ∙ 119891119909
where fx = strain incompatibility factor
fx = 010 for drilled shafts augered cast-in-place and bored piles
fx = 020 for driven low displacement piles (opened-ended steel pipe and H-piles)
fx = 030 for driven high-displacement piles (ie solid piles PSC and closed-ended pipe)
C-12
For undrained loading of circular piles in compression no reduction of end-bearing is necessary
therefore
119902119887 = 933 middot 119904119906 C11
C12 Direct CPT Methods for Pile Capacity
In the direct CPT method the penetrometer readings are scaled directly via specified algorithms to
obtain the pile unit side friction and end-bearing as depicted in Figure C4 As many as 40 different
direct CPT methods have been developed over the past five decades as summarized by Niazi amp Mayne
(2013) Starting circa 1970 many of these early methods relied on hand-recorded data from field
mechanical CPTs where only qc data were obtained at 20 cm intervals or later with mechanical readings
of both qc and fs using special sets of inner and outer rods that recorded vertical load for tip and loads
for tip plus sleeve in alternating increments Also early pile load test data were often obtained from
top-down measurements of load-displacement
C-13
Undrained qb = Nc su
Qside = S (fp DAs)
QTotal = Qs + Qb - Wp
fp = cmck Ko svorsquo tanrsquo
Qbase = qb Ab
Drained qb = Nq svorsquo
Method Two Rational or ldquoIndirectrdquo Method
Method OneldquoDirectrdquo CPT Method
(Scaled Pile)
fp = fctn (soil type pile
type qt fs and u2)
qb = fctn (soil type qt-u2)
qb = unit end bearing
unit sidefriction fp
OCR su Ko t DR rsquo
AXIAL PILE CAPACITY FROM CPT READINGS
Figure C4 Concept of Direct versus Rational Method for CPT evaluation of axial pile capacity
Beginning in the mid-1990s as the modern electric piezocone (CPTu) was implemented the newer
equipment offered improved data and better resolution because of the additional reading of porewater
pressures and correction of raw measured qc to total resistance qt In addition the use of electronic
digital data collection and field computers proved superior in field measurements and recordings As a
consequence several reliable CPT methods for axial pile capacity have been developed Moreover
parallel improvements in full-scale pile load testing occurred and now include modern strain gage
instrumentation digital data recording and automated testing procedures Also the testing can be
single direction (compression or tension) or bi-directional as in the Osterberg cell
C-14
Table C3 provides a selection of recent direct CPT methods that have become available over the
past two decades A number of these (namely ICP NGI UWA and Fugro) were funded by the
offshore industry because of the growth of oil amp gas reserves and windfarm installations thus
necessitating increased concerns on risk probability and reliability in the site investigations for
offshore platforms and design of large driven monopile foundations These direct CPT methods
are often statistically based on large datasets compiled from full-scale load tests made
worldwide (eg Schneider et al 2008)
Table C3 Selection of Direct CPT Methods for Axial Pile Capacity
Method Pile Types Soil Types References CPT
data
Additional
parameters
needed
Unicone Method all types Sands silts
clays
Eslami and
Fellenius
(1997)
Fellenius
(2009)
qt fs u2
KTRI = Kajima
Technical
Research
Institute
all types Sands
mixed clays
Takesue et al
(1998)
fs and u2 No guidance given
on end bearing
resistance
Imperial College
Procedure (ICP)
OE and CE Sands Chow et al
(1997 PhD)
Jardine et al
(2005)
qt Interface friction
angle (δ) from
ring shear tests
Correlated to
mean grain size
(D50)
C-15
OE and CE Clays Jardine et al
(2005)
qt Interface friction
angle (δ)
correlated to PI
Table C3 Continued
NGI Method
(Norwegian
Geotechnical
Institute)
OE and CE Sands Clausen et al
(2005)
qt DR from qt
Clays a Alpha
method
(Karlsrud et al
2005 2012)
b Beta
method
(Karlsrud
2012)
qt for su
qt for
OCR
PI = plasticity
index ()
PI = plasticity
index ()
Fugro Method OE and CE Sands Kolk et al
(2005)
qt
Clays Van Dijk and
Kolk (2011)
qt
OE and CE Sands Lehane et al
(2005)
qt IFR = infilling ratio
related to amount
of plugging
C-16
Purdue LFRD Drilled
shafts
Sands Basu amp
Salgado
(2012)
qt LRFD = load
resistance
factored design
Enhanced Various Sands silts Niazi and qt u2 Based on 330 load
Unicone Method clays and Mayne (2015 and fs tests including
mixed soils 2016) bored augered
jacked and driven
piles
UWA (Univ of
Western
Ra = pile
roughness
Australia) Clays Lehane et al
(2012)
qt Interface friction
angle (δ) from
ring shear tests
and also method
without δ
HKU Method
(Hong Kong
Univ)
OE and CE
(end
resistance
only)
Sands Yu and Yang
(2012)
qt End bearing only
PLR = plug length
ratio
Note PLR can be
estimated from
OE inner diam
Table C3 Continued
Note previously called Marine Technical Directorate (MTD)
C-17
Many of the offshore CPT methods relate to driven piles either in sand or clay as that is
common deep foundation for those purposes Of particular interest are the Unicone and
Modified Unicone Methods since they use all three readings of the piezocone (qt fs and u2)
and address a variety of pile foundation types
C13 Modified UniCone Method
The modified UniCone Method was developed based on a total 330 pile load tests which were
associated with SCPTu data during their site investigations (Niazi and Mayne 2015 2016) This
represents a threefold increase over the original UniCone database that was built upon data
from 106 pile load tests (Eslami amp Fellenius 1997)
For the original UniCone algorithms use is made of the effective cone resistance (qE)
119902119864 = 119902119905 minus 1199062 C12
and a chart of qE vs fs provided an approximate soil classification in five distinct groups as shown by
Figure C5 Later in the modified approach a better delineation of the larger dataset gave soil
subclassifications as indicated by Figure C6
C-18
Figure C5 Soil behavior type using CPT via original UniCone
Figure C6 Soil behavior type using CPT via Modified UniCone
C-19
In the modified approach the 9-zone normalized soil behavioral type (SBTn) is ascertained using CPT
data in conjunction with the Robertson (2009) charts as shown in Figure C7 As discussed in Appendix
A and B the SBTn method uses the normalized cone resistance (Qtn) normalized sleeve friction (Fr())
and CPT material index Ic This permits a much wider range in the calculated pile side friction because fp
is related as a continuous curve with Ic rather than only 5 values that are assigned in the original
scheme
The pile unit side friction (fp) is obtained from qE and the CPT material index Ic using the following
expression at each elevation along the sides of the pile
10(0732 middot 119868119888 minus 3605) fp middot middot C13 = 119902119864 middot 120579119875119879 middot 120579119879119862 120579119877119860119879119864
C-20
Figure C7 Nine-zone soil behavioral soil type using normalized piezocone parameters
(after Robertson 2009 Mayne 2014)
where θPT = coefficient for pile type (θPT = 084 for bored piles 102 for jacked piles 113 for driven
piles) θTC = coefficient for loading direction (θTC = 111 for compression and 085 for tension) and θRATE
= rate coefficient applied to soils in SBT zones 1 through 7 (θRATE = 109 for constant rate of penetration
tests and 097 for maintained load tests) Since CPT provides data at regular intervals of 2 cm to 5 cm
along the sides of the pile the average fp from z = 0 to z = L can be used directly in Equation 1 to obtain
the shaft capacity Qs
C-21
The pile end bearing resistance is obtained from
middot 10(0325middot 119868119888minus1218) 119902119887 = 119902119864 C14
where qE is averaged in the vicinity of the pile tip Figure C8 shows the Modified Unicone Method in
graphical format For sensitive clays of zone 1 please see additional discussions by Niazi amp Mayne
(2016)
Figure C8 Summary of Modified UniCone Method for direct CPT assessment of axial pile
capacity
C-22
C14 Axial Pile Displacement
The movement of pile foundations can be assessed using elastic continuum theory (Poulos amp
Davis 1980 Randolph 2003 Mayne amp Niazi 2017) These relationships have been developed
using finite element analyses boundary elements and analytical closed-form solutions For the
latter approach the top displacement of a rigid pile subjected to a vertical force is shown in
Figure C9 This gives the movement at the top of a pile subjected to either compression or
tension (uplift) loading Also the percentage of axial load transferred to the pile toe is
determined For piles extending through various soil layers the elastic solution can be
implemented by using a set of stacked pile segments each with its own stiffness as
represented by a soil Youngs modulus
For pile groups the use of computer software is recommended Several available programs can handle
pile groups under axial and lateral moment loading such as DEFPIG (Univ Sydney) and PIGLET (Univ
Western Australia) A full listing of pile foundation software is given at the Geotechnical amp
GeoEnvironmental Service Directory wwwggsdcom
Ps
= Sh
aft
Load
sL
t
tEd
IPw
Load Transfer to Base
)]1)((5ln[
)(
)1(1
1
2 vdL
dLFI
E
ECT
211
IF
P
P CT
t
b
Base Load = Pb
Randolph Elastic Solution
Ground Surface
Top Load Pt = Ps + Pb
Rigid Pile Response
wt = displacementd = diameter
L = length
Es = Elastic Soil Modulus
Es0 (at z = 0) surface
EsM (at z = L2)mid-length
EsL (at z = L)full length
Load Transfer to Shaft
E = EsMEsL = Gibson parameterE = 1 for homogeneous case (constant E)E = 05 for pure Gibson case (Es0 = 0)
where FCT = load direction (= 1 compression 0 tension)
21
IF
P
P CT
t
b
z = depth
= Poissons ratio
Tension
Compression
C-23
Figure C9 Elastic continuum solution for axial pile response under compression and tension
loading
C-24
C15 Acknowledgements
The author appreciates the help of Fernando Illingworth of PonsEcuador Meena Viswanth of
GeoSyntecAtlanta and Dr Marco Uzielli of GeoRiskItaly in helping to collect and process the data
Funding support was provided by David Woeller of ConeTec Richmond BC
C-25
- Structure Bookmarks
-
- CONE PENETRATION TEST DESIGN GUIDE FOR STATE GEOTECHNICAL ENGINEERS
- TABLE OF CONTENTS
- LIST OF TABLES
- CHAPTER 1 INTRODUCTION
- CHAPTER 2 DIRECT CPT METHOD FOR SOIL CHARACTERIZATION
- CHAPTER 3 DIRECT CPT METHOD FOR SHALLOW FOUNDATIONS
- CHAPTER 4 DIRECT CPT METHOD FOR DEEP FOUNDATIONS
- REFERENCES
- APPENDIX A
- APPENDIX B
- APPENDIX C
-
1
2
3
4
5
6
7
8
9
10
11
12
13
14
15
16
17
18
19
20
21
22
23
24
LIST OF FIGURES
Figure Geoparameters determined from CPT 1
Figure Conventional method for shallow foundation design compared to direct CPT method 2
Figure Direct CPT evaluation of axial pile capacity 3
Figure Soil unit weight from CPT sleeve friction 6
Figure Approximate ldquorules of thumbrdquo method using CPT sounding from Wakota Bridge MN 7
Figure SBT zones using CPT Ic 9
Figure Yield stress exponent compared to CPT material index 11
Figure k vs dissipation time for 50 consolidation (Mayne 2017) 15
Figure CPT data from Benton County Minnesota for example problem 1 16
Figure Soil layers using ldquorules of thumbrdquo17
Figure Soil layer 1 using SBT method 23
Figure Soil layer 2 using SBT method 24
Figure Soil layer 3 using SBT method 25
Figure Soil layer 4 using SBT method 26
Figure CPT data from Minnesota for example problem 2 34
Figure Soil layers using ldquorules of thumbrdquo36
Figure Soil layer 1 using SBT method 42
Figure Soil layer 2 using SBT method 43
Figure Soil layer 3 using SBT method 44
Figure Soil layer 4 using SBT method 45
Figure Dissipation t50 data 53
Figure Direct CPT method for shallow foundations56
Figure Direct CPT method introduction 58
Figure Differentiation of porewater pressure measurement locations (Lunne et al 1997) 58
Figure 25 Foundation soil formation parameter hs versus CPT material index Ic (Mayne 2017) 60
Figure 26 Diagram of footing profiles for Example 362
Figure 27 CPT data from Northern Minnesota 63
Figure 28 Schematic of foundation design associated with CPT data 64
Figure 29 Soil type for example problem 367
Figure 30 Direct CPT evaluation of axial pile capacity 70
Figure 31 UniCone Method soil behavior type using CPT (Mayne 2017) 71
Figure 32 Modified UniCone Method soil behavior type using CPT (Mayne 2017) 71
Figure 33 Deep foundation end bearing pile diagram74
Figure 34 CPT data from Minnesota for example problem 5 75
Figure 35 Soil layers using ldquorules of thumbrdquo for pile capacity example using direct CPT method 76
Figure 36 Soil layer 1 using SBT method 81
Figure 37 Soil layer 2 using SBT method 82
Figure 38 Soil layer 3 using SBT method 83
Figure 39 Soil layering compared to pilings 85
LIST OF TABLES
Table 1 Geoparameters calculated directly from CPT 4
Table 2 Minor Geoparameters 5
Table 3 Soil type compared to exponent m 10
Table 4 Geoparameters evaluated for Case Example 117
Table 5 Geoparameters 35
Table 6 Bearing capacity defined by soil type61
Table 7 SCPT Results 62
CHAPTER 1 INTRODUCTION
The purpose of this manual is to provide an analysis of soil characterization shallow footings and deep
foundations using direct cone penetration testing (CPT) methods Geotechnical site characterization is
important for evaluating soil parameters that will be used in the analysis and design of foundations
retaining walls embankments and situations involving slope stability Common practice is to determine
these soil parameters through conventional lab and in-situ testing An alternative method uses CPT
readings of cone tip resistance (qt) sleeve friction ( fs) and pore pressure (u2) directly to determine these
parameters such as unit weight effective friction angle undrained shear strength and many others
(Figure 1) Direct CPT methods are provided for these parameters which will be used in the designs for
shallow and deep foundations
Figure 1 Geoparameters determined from CPT
Designing shallow foundations is typically done in a two-part process determining the bearing capacity
and expected settlement (commonly referred to as displacement) of the soil to approximate the
required size and shape of a foundation The older traditional methods are no longer required with
many approaches existing for using CPT directly in the design of shallow foundations Results from these
methods can provide a direct assessment of bearing capacity andor settlement A specific approach to
the direct method has been recommended and tested using a database of 166 full-scale field load tests
(Figure 2) A one-part process is used to scale the measured cone penetrometer readings (ie
measured cone tip resistance sleeve friction and pore water pressure) to obtain the bearing capacity
and settlement of the soil A step-by-step procedure has been created to transition from the CPT data to
bearing capacity with settlement accounted for The steps consist of estimating a design footing width
1
and length while using the process in the Soil Characterization section to determine soil parameters
directly from the CPT data
Figure 2 Conventional method for shallow foundation design compared to direct CPT method
There are upwards of 40 different direct CPT methods that have been developed over the past five
decades to determine the axial compression capacity of a piling foundation Earlier direct CPT methods
relied on hand-recorded information where mechanical-type CPT cone tip resistance data would be
collected at 20 cm intervals
Herein the method recommended for deep foundation design is the Modified UniCone method which
uses all three readings of the modern electronic piezocone penetrometer (CPTu) while addressing a
variety of pile foundation types (Figure 3) The modified UniCone Method is based on a total of 330 pile
load tests (three times the original UniCone database) that were associated with SCPTu data Two
computer software programs have been recommended for analyzing pile movements andor
settlements
2
Figure 3 Direct CPT evaluation of axial pile capacity
3
Symbol Parameter
γt Soil total unit weight
Ic CPT material index
SBT Soil behavior type (SBT)
ʹ σp Preconsolidation stress
YSR Yield stress ratio
ϕʹ Effective friction angle
Ko Lateral stress coefficient
su Undrained shear strength
Dʹ Constrained modulus
Eʹ Drained Youngrsquos modulus
CHAPTER 2 DIRECT CPT METHOD FOR SOIL
CHARACTERIZATION
21 INTRODUCTION
A direct CPT method for determining the value of each of various geoparameters shown in Table 1 is
provided The parameter Ic is used in the derivation of several sequential parameters This section of
the guide may be referred to as these parameters are used in calculations throughout the shallow
foundations and deep foundations sections
Table 1 Geoparameters calculated directly from CPT
4
Kʹ Bulk modulus
ks Subgrade reaction modulus
MR Resilient modulus
Gmax Small-strain shear modulus
k Coefficient of permeability
cv Coefficient of consolidation
Symbol Parameter Equation
ρt Mass Density ρt = γtga
where ga = 98 ms2
σvo Total Stress σvo asymp Σ (γti middot Δzi)
c Effective cohesion ʹ Empirical c asymp 003σp
In clays c asymp 01cu
Other minor parameters can be used in calculations of the geoparameters in Table 1 These minor parameters are provided in Table 2
Table 2 Minor Geoparameters
5
22 SOIL UNIT WEIGHT
The total soil unit weight can be estimated from CPT sleeve friction resistance as shown in Figure 4
(Mayne 2014) This method is not applicable to organic clays diatomaceous soils peats or sensitive
soils
Figure 4 Soil unit weight from CPT sleeve friction
Use Equation 1 to calculate the soils total unit weight
1 120574119905 = 120574119908 ∙ [122 + 015 ∙ ln (100 ∙ 119891119904 + 001)]
120590119886119905119898
Since soil unit weight is required for determining most geoparameters (including Soil Behavior Type)
estimating the soil type of the layers using the ldquorules of thumbrdquo method is a good first step before
determining more precise layering (Figure 5) Once a unit weight is determined for each noticeable layer
change these results can be used in later calculations such as ldquoCPT Material Indexrdquo To use the ldquorules
of thumbrdquo method some helpful guidelines are to assume sands are identified when qt gt 725 psi and u2
asymp uo while the presence of intact clays are prevalent when qt lt 725 psi and u2 gt uo The magnitude of
porewater pressures help to indicate intact clays such as soft (u2 asymp 2middotuo) firm (u2 asymp 4middotuo) stiff (u2 asymp 8middotuo)
and hard (u2 asymp 20middotuo) Fissured overconsolidated clays tend to have negative u2 values such that u2 lt 0
6
Figure 5 Approximate ldquorules of thumbrdquo method using CPT sounding from Wakota Bridge MN
23 CPT MATERIAL INDEX
The development of the CPT material index Ic has improved the initial classification of soil types and
calculation of soil parameters as shown in Table 1 To calculate Ic follow steps 1 and 2
231 Step 1 Normalized Sleeve Friction
Calculate Fr using Equation 2 if not provided with the CPT data gathered
2
7
119891119904 119865119903() = 100 ∙
(119902119905minus120590119907119900)
232 Step 2 Iteration
Iterate using Equations 3-5 to determine Ic by initially using n = 1 to calculate a starting value of Ic The
exponent n is soil-type dependent n = 1 (clays) n asymp 075 (silts) and n asymp 05 (sands) Iteration converges
quickly which is generally after the 3rd cycle
3 (119954119957minus120648119959119952)120648119938119957119950 = 119951 119928119957119951 prime (120648119959119952120648119938119957119950)
4
prime 120590119907119900 119899 = 0381 ∙ 119868119888 + 005 ( ) minus 015 120590119886119905119898
119899 le 10
5 119920119940 = radic[120785 120786120789 minus 119949119952119944119928119957119951]120784 + [120783 120784120784 + 119949119952119944119917119955]120784
24 SOIL BEHAVIOR TYPE (SBT)
Typically soil samples are not taken when using CPT The soil types are then inferred from the qt fs and
u2 readings To determine the types of soil from CPT data follow steps 1 and 2
241 Step 1
To determine the soil layers from CPT results calculate Ic by following the steps under section ldquoCPT
Material Indexrdquo After Ic has been determined through all specified depths use Figure 6 to classify the
type of soil by comparing each Ic value to normalized CPT readings (Fr and Qtn) from Equations 2 and 3
8
Figure 6 SBT zones using CPT Ic
242 Step 2
To determine if any of the soil layers contain ldquosensitive clays and siltsrdquo from zone 1 or ldquovery stiff
overconsolidated (OC) soilrdquo from zones 8 and 9 use Equations 6 and 7 If any soil layers are found within
zone 1 by Equation 6 then caution should be taken as these clays are prone to instability collapses and
difficulties in construction performance Very stiff OC sands to clayey sands of zone 8 (15 lt Fr lt 45)
and very stiff OC clays to silts of zone 9 (Fr gt 45) can be identified by Equation 7
6 Equation 6 errata This is an exponential expression (see Figure 5)
119876119905119899 lt 12119890119909119901(minus14 ∙ 119865119903)
7 119876119905119899 gt 0005(119865119903minus1)minus00003(119865119903minus1)2minus0002
1
Errata See terms and coefficients in Figure 5 above (numbers cannot be rounded off)
After each CPT reading has been assigned a zone from Figure 14 a visual representation can be made to
show the predominant layers by soil types (Figure A5)
25 EFFECTIVE STRESS FRICTION ANGLE
The effective friction angle (ϕ) is used to govern the strength for sands and clays where Equation 8 is
used for sands and Equation 9 is used for clays The value of ϕ for sands is derived from Ic so before ϕ
can be calculated refer to the iteration of Qtn under the ldquoCPT Material Indexrdquo section Once the values
9
Soil Type m
Fissured clays 11
Intact clays 10
Sensitive clays 09
Silt mixtures 085
Silty sands 080
Clean sands 072
Note m may be higher than 11 in fissured clays
of Ic and Qtn are calculated the type of soil can be determined from the section ldquoSoil Behavior Type
SBTrdquo The type of soil will dictate which equation to use for ϕ
Sands
8 120601prime(deg) = 176deg + 110deg log(119876119905119899)
Clays
9 119902119905minus120590119907119900 120601prime(deg) = 295deg ∙ 119861119902
0121 ∙ [0256 + 0336 ∙ 119861119902 + log ( )] prime 120590119907119900
Where = 119861119902 (1199062 minus 119906119900)(119902119905 minus 120590119907119900)
26 STRESS HISTORY
Determining the stress history can be characterized by an apparent yield stress ratio of the form
10 prime 120590119901
119884119878119877 = prime 120590119907119900
Where σp is defined as the preconsolidation stress or effective yield stress (Equation 10) The YSR is the
same equation as the more common overconsolidation ratio (OCR) but is now generalized to
accommodate mechanisms of preconsolidation such as ageing desiccation repeated cycles of wetting-
drying repeated freeze thaw cycles and other factors
11 prime prime 120590119901 = 120590119910 = 033(119902119905 minus 120590119907119900)119898prime
Where σp σy σvo and qt have units of kPa and the value of m depends on soil type with typical values shown in Table 3
Table 3 Soil type compared to exponent m
10
The value of m for non-fissured soils and inorganic clays and silts is derived from Ic (Figure 7) so before
m can be calculated (Equation 12) refer to the iteration of Ic under the ldquoCPT Material Indexrdquo section
Determine Ic for all soil layers before calculating m
12 028
119898prime = 1 minus 25 119868119888 1+( frasl ) 265
Once m is known for all soil layers the YSR can be determined
Figure 7 Yield stress exponent compared to CPT material index
A limiting value of YSR can be reached for clays and sands It can be calculated using Equation 13
13 (1 sin(emptyprime))
(1+sin(emptyprime)) 119884119878119877119897119894119898119894119905 = [ ]
(1minussin(emptyprime))2
27 LATERAL STRESS COEFFICIENT
The lateral stress coefficient Ko = σhoσvo commonly referred to as the at-rest condition is used to
represent the horizontal geostatic state of soil stress Ko can be calculated using Equation 14 but
11
14
119870119900 = (1 minus sin(emptyprime)) ∙ 119884119878119877sin( emptyprime)
sections such as ldquoStress Historyrdquo and ldquoEffective Stress Friction Anglerdquo will need to be referred to for
determining parameters ϕ and YSR
A maximum value for Ko can be determined by Equation 15
15 (1+sin(emptyprime))
119870119900119898119886119909 = = 1199051198861198992(45deg + emptyprime2) (1minussin(emptyprime))
28 UNDRAINED SHEAR STRENGTH
Loading on soils can result in fully drained partially drained or fully undrained conditions Sands
typically produce drained cases due to their high permeability but exceptions may occur in loose sands
during fast loading where the water does not have sufficient time to dissipate Clays exhibit low
permeability and thus often result in undrained loading cases when a load is applied quickly For soft-
firm clays the undrained shear strength (su) can be determined from CPT via Equation 16 where the
value of the bearing factor Nkt can be taken as 12
16 119902119905minus120590119907119900 119904119906 =
119873119896119905
In the case of remolded undrained shear strength from CPT su asymp fs
29 GROUND STIFFNESS AND SOIL MODULI
Determining the grounds stiffness can be measured from geoparameters such as the constrained
modulus (D) drained Youngrsquos modulus (E) bulk modulus (K) subgrade reaction modulus (ks) resilient
modulus (MR) and small-strain shear modulus (Gmax)
17 119863 prime asymp 5 ∙ (119902119905 minus 120590119907119900)
18
12
119864 prime = 119863 prime
11
19 119864prime
119870prime = [3∙(1minus2119907prime)]
The subgrade modulus (ks) is a combination of soil-structural properties which creates a parameter that
depends on the ground stiffness and the size of the loaded element
20 119864prime
119896119904 = [119889∙(1minus1199072)]
The resilient modulus MR applies to pavement analysis and design and can be calculated using Equation
21 where qt and fs are in MPa
21 053 119872119877 = (146119902119905 + 135511989111990414 + 236)244
The small strain shear modulus Gmax is a representation of the initial stiffness of all soils and rocks
Graphically it is the beginning portion of all stress-strain-strength curves for geomaterials Use Equation
22 to determine Gmax
22 119866119898119886119909 = 120588119905 ∙ 1198811199042
Where Vs = shear wave velocity (Equation 23) as measured by seismic cone penetration tests (SCPT) If
only cone penetration tests (CPT) or piezocone (CPTu) data are available the shear wave velocity may
be estimated from
23 03 119891119904 119881119904 (119898frasl119904) = [101 ∙ log(119902119905) minus 114]167 ∙ (100 ∙ )
119902119905
Where qt and fs have units of (kPa)
210 COEFFICIENT OF CONSOLIDATION
The coefficient of consolidation (cv) controls the rate that foundation and embankment settlements
occur By using results of CPT dissipation tests that measure the change in u2 readings over time the
value of cv can be determined Using Equation 24 cv can be determined based on piezocone dissipation
13
curves The equation below requires an estimate of the in-situ rigidity index (IR) of the soil If results of
SCPTU are available then IR may be determined from Gmax and qt per equation A38
24 0030∙(119886119888)2∙(119868119877)075
119888119907 = 11990550
Where ac = penetrometer radius (178 cm for 10-cm2 cone 220 cm for 15-cm2 cone) t50 = time to reach 50 dissipation IR = Gsu = undrained rigidity index G can be determined from the ldquoGround Stiffness and Soil Modulirdquo section
211 HYDRAULIC CONDUCTIVITY
The hydraulic conductivity also known as the coefficient of permeability (k) expresses the flow
characteristics of soils and has units of cms or feetday One method of calculation would be to use
Equation 25 where cv would need to be determined from ldquoCoefficient of Consolidationrdquo and D would
need to be determined from ldquoGround Stiffness and Soil Modulirdquo
25 119888119907∙120574119908 119896 =
119863prime
An alternative approach developed for soft normally-consolidated soils (Figure 8) is shown in Equation
26 where t50 (sec) values are used directly in assessing k in (cms)
26
14
125
119896 asymp ( 1
) 251∙11990550
Figure 8 k vs dissipation time for 50 consolidation (Mayne 2017)
212 EXAMPLE PROBLEMS
2121 Example 1 Direct CPT Methods for Geoparameters on Sands
Several geoparameters need to be determined based on the given CPT data collected for the South
Abutment of a bridge in Benton County MN (Figure 9) The groundwater table (GWT) was measured at
17 feet Determine all the geoparameters found in Table 4 at depths of 0 feet to 30 feet All sand layers
can be assumed ldquodrainedrdquo with ν = 02
15
Figure 9 CPT data from Benton County Minnesota for example problem 1
16
Table 4 Geoparameters evaluated for Case Example 1
Symbol Parameter Symbol Parameter
γt Soil total unit weight Ko Lateral stress coefficient
Ic CPT material index Dʹ Constrained modulus
SBT Soil behavior type (SBT) Eʹ Drained Youngrsquos modulus ʹ σp Preconsolidation stress Kʹ Bulk modulus
YSR Yield stress ratio MR Resilient modulus
ϕʹ Effective friction angle Gmax Small-strain shear modulus
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 10 Soil layers using ldquorules of thumbrdquo
γw = 6224 pcf
17
Layer 1
From Figure 10 fs = 17 psi taken as a representative value of the layer
17 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1204 pcf
145 psi
Layer 2
From Figure 10 fs = 17 psi taken as a representative value of the layer
17psi
γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1204 pcf 145 psi
Layer 3
From Figure 10 fs = 12 psi taken as a representative value of the layer
12 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 4
From Figure 10 fs = 7 psi taken as a representative value of the layer
CPT Material Index
7 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1121 pcf
145 psi
Layer 1
From Figure 10 fs = 17 psi and qc = 3500 psi
18
qt = qc + u2 ∙ (1 minus a) = 3500 psi + 3 psi ∙ (1 minus 08) = 3501 psi
= γt ∙ 6 feet = 1204 pcf ∙ 6 feet = 7225 psf = 50 psi σvo
fs 17 psi Fr() = 100 ∙ = 100 ∙ = 049
(qt minus σvo) (3501 psi minus 50 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 5 psi minus 0 psi = 50 psi
(qt minus σvo)σatm (3501 psi minus 50 psi )145 psi = = Qtn prime )n (σvoσatm (50 psi145 psi)n
prime σvo 50 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(049)]2
Qtn = 35329 n = 036 lt 10 Ic = 13
Layer 2
From Figure 10 fs = 17 psi and qc = 3500 psi
19
qt = qc + u2 ∙ (1 minus a) = 3500 psi + 0 psi ∙ (1 minus 08) = 3500 psi
= 7225 psf + γt ∙ 6 feet = 7225 psf + 1204 pcf ∙ 6 feet = 1445 psf = 100 psi σvo
fs 17 psi Fr() = 100 ∙ = 100 ∙ = 049
(qt minus σvo) (3500 psi minus 100 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 100 psi minus 0 psi = 100 psi
(qt minus σvo)σatm (3500 psi minus 100 psi )145 psi = = Qtn prime )n (σvoσatm (100 psi145 psi)n
prime σvo 100 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
47 minus log(Qtn)]2 + [122 + log(049)]2
Ic = radic[3
Qtn = 27947 n = 041 lt 10 Ic = 14
Layer 3
From Figure 10 fs = 12 psi and qc = 1500 psi
20
qt = qc + u2 ∙ (1 minus a) = 1500 psi + 0 psi ∙ (1 minus 08) = 1500 psi
= 1445 psf + γt ∙ 11 feet = 1445 psf + 1172 pcf ∙ 11 feet = 2734 psf = 190 psi σvo
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 081
(qt minus σvo) (1500 psi minus 190 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = γwater ∙ (z minus 119911119908) = 6224 pcf ∙ (23 feet minus 17 feet) = 3734 psf = 99 psi
prime σvo = σvo minus uo = 190 psi minus 99 psi = 164 psi
(qt minus σvo)σatm (1500 psi minus 190 psi )145 psi = = Qtn prime )n (σvoσatm (164 psi145 psi)n
prime σvo 164 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(081)]2
Qtn = 947 n = 06 lt 10 Ic = 19
21
Layer 4
From Figure 10 fs = 7 psi and qc = 1200 psi
qt = qc + u2 ∙ (1 minus a) = 1200 psi + 0 psi ∙ (1 minus 08) = 1200 psi
= 2734 psf + γt ∙ 11 feet = 2734 psf + 1121 pcf ∙ 7 feet = 3519 psf = 244 psi σvo
fs 7 psi Fr() = 100 ∙ = 100 ∙ = 060
(qt minus σvo) (1200 psi minus 244 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = γwater ∙ (z minus 119911119908) = 6224 pcf ∙ (30 feet minus 17 feet) = 8091 psf = 130 psi
prime σvo = σvo minus uo = 244 psi minus 130 psi = 188 psi
(qt minus σvo)σatm (1200 psi minus 244 psi )145 psi = = Qtn prime )n (σvoσatm (188 psi145 psi)n
prime σvo 188 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(060)]2
22
Qtn = 686 n = 064 lt 10 Ic = 19
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic Qtn and Fr the first layer is defined as a ldquoDrained Gravelly Sandrdquo from
Figure 11
Figure 11 Soil layer 1 using SBT method
23
Layer 2
Based on values of Ic Qtn and Fr the second layer is defined as a ldquoDrained Sandrdquo from Figure
12
Figure 12 Soil layer 2 using SBT method
24
Layer 3
Based on values of Ic Qtn and Fr the third layer is defined as a ldquoDrained Sandrdquo from Figure 13
Figure 13 Soil layer 3 using SBT method
25
Layer 4
Based on values of Ic Qtn and Fr the fourth layer is defined as a ldquoDrained Sandrdquo from Figure 14
Figure 14 Soil layer 4 using SBT method
Effective Stress Friction Angle
Layer 1
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(3533) = 456deg
Layer 2
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(2795) = 445deg
26
Layer 3
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(947) = 393deg
Layer 4
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(686) = 378deg
Stress History
Layer 1
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (13frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(24139 kPa minus 345 kPa)072 = 4717 kPa
= 684 psi
prime σp 684 psi YSR = = = 136
primeσvo 50 psi
Layer 2
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + (14 1 + ( frasl ) frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(34132 kPa minus 689 kPa)072 = 605 kPa
= 877 psi
prime σp 877 psi YSR = = = 87
primeσvo 100 psi
27
Layer 3
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + (19 1 + ( frasl ) frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(10342 kPa minus 131 kPa)072 = 254 kPa
= 368 psi
prime σp 368 psi YSR = = = 22
primeσvo 164 psi
Layer 4
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (19frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(8274 kPa minus 168 kPa)072 = 215 kPa = 312 psi
prime σp 312 psi YSR = = = 17
primeσvo 188 psi
Lateral Stress Coefficient
Layer 1
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(456deg)) ∙ 136sin( 456deg) = 18
Layer 2
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(445deg)) ∙ 87sin( 445deg) = 14
28
Dprime 17480 psi
Eprime = = = 15890 psi 11 11
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(393deg)) ∙ 22sin( 393deg) = 06
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(378deg)) ∙ 17sin( 378deg) = 05
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3501 psi minus 50 psi) = 17480 psi
Eprime 15890 psi
Kprime = = = 8828 psi [3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Layer 3
Layer 4
Ground Stiffness and Soil Moduli
Layer 1
Values of qt and fs need to be in MPa
MR = (146(241 MPa)053 + 1355(012 MPa)14 + 236)244 = 3416 MPa = 49575 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1172 kPa m Vs (mfrasls) = [101 ∙ log(24139 kPa) minus 114]167 ∙ (100 ∙ ) = 2745
24139 kPa s
Vs = 9012 fts
29
γt 1204 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 37 ∙ (9012 ) = 304 ∙ 106psf = 21000 psi s
Layer 2
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3500 psi minus 100 psi) = 17480 psi
Dprime 17480 psi Eprime = = = 15890 psi
11 11
Eprime 15890 psi Kprime = = = 8828 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(241 MPa)053 + 1355(012 MPa)14 + 236)244 = 3415 MPa = 49562 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1172 kPa m Vs (mfrasls) = [101 ∙ log(24133 kPa) minus 114]167 ∙ (100 ∙ ) = 2745
24133 kPa s
Vs = 9012 fts
γt 1204 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
30
2 ft Gmax = ρt ∙ Vs
2 = 37 ∙ (9012 ) = 304 ∙ 106psf = 21000psi s
Layer 3
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (1500 psi minus 190 psi) = 7405 psi
Dprime 7405 psi Eprime = = = 6732 psi
11 11
Eprime 6732 psi Kprime = = = 3740 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(103 MPa)053 + 1355(008 MPa)14 + 236)244 = 1506 MPa = 21854 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 827 kPa m Vs (mfrasls) = [101 ∙ log(10343 kPa) minus 114]167 ∙ (100 ∙ ) = 2611
10343 kPa s
Vs = 8566 fts
γt 1172 pcf slug ρt = = = 36
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 36 ∙ (8566 ) = 27 ∙ 106psf = 19000 psi s
Layer 4
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (1200 psi minus 244 psi) = 5878 psi
31
Dprime 5878 psi Eprime = = = 5343 psi
11 11
Eprime 5343 psi Kprime = = = 2969 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(83 MPa)053 + 1355(005 MPa)14 + 236)244 = 165 MPa = 16901 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 483 kPa m Vs (mfrasls) = [101 ∙ log(8274 kPa) minus 114]167 ∙ (100 ∙ ) = 2243
8274 kPa s
Vs = 7360 fts
γt 1121 pcf slug ρt = = = 35
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 35 ∙ (7360 ) = 389 ∙ 106psf = 13000 psi s
32
2122 Example 2 Direct CPT Methods for Geoparameters on Clay
Several geoparameters need to be determined based on the given CPT data collected for the South
Abutment (Figure 15) The groundwater table (GWT) was measured at 60 feet Determine all the
geoparameters found in Table 5 at depths of 0 feet to 42 feet All sand layers can be assumed ldquodrainedrdquo
ν = 02 All clay layers can assume ν = 049
33
Figure 15 CPT data from Minnesota for example problem 2
34
Table 5 Geoparameters
Symbol Parameter Symbol Parameter
γt Soil total unit weight Eʹ Drained Youngrsquos modulus
Ic CPT material index Kʹ Bulk modulus
SBT Soil behavior type (SBT) MR Resilient modulus
ʹ σp Preconsolidation stress Gmax Small-strain shear modulus
YSR Yield stress ratio su Undrained shear strength
ϕʹ Effective friction angle cv Coefficient of consolidation
Ko Lateral stress coefficient k Hydraulic conductivity
Dʹ Constrained modulus
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
35
Figure 16 Soil layers using ldquorules of thumbrdquo
Unit weight of water γw = 6224 pcf
Layer 1
From Figure 16 fs = 13 psi taken as a representative value of the layer
13 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1179 pcf
145 psi
Layer 2
From Figure 16 fs = 12 psi taken as a representative value of the layer
12 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 3
From Figure 16 fs = 20 psi taken as a representative value of the layer
36
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
Layer 4
From Figure 16 fs = 2 psi taken as a representative value of the layer
2 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1004 pcf
145 psi
CPT Material Index
Layer 1
From Figure 10 fs = 13 psi and qc = 3000 psi
qt = qc + u2 ∙ (1 minus a) = 3000 psi + 0 psi ∙ (1 minus 08) = 3000 psi
σvo = γt ∙ 2 feet = 1179 pcf ∙ 2 feet = 235 psf = 16 psi
fs 13 psi Fr() = 100 ∙ = 100 ∙ = 043
(qt minus σvo) (3000 psi minus 16 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 16 psi minus 0 psi = 16 psi
(qt minus σvo)σatm (3000 psi minus 16 psi )145 psi = = Qtn prime )n (16 psi145 psi)n (σvoσatm
prime σvo 16 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(043)]2
37
Qtn = 4126 n = 032 lt 10 Ic = 121 therefore sand
Layer 2
From Figure 10 fs = 12 psi and qc = 250 psi
qt = qc + u2 ∙ (1 minus a) = 250 psi + 10 psi ∙ (1 minus 08) = 252 psi
σvo = 235 psf + γt ∙ 30 feet = 235 psf + 1172 pcf ∙ 30 feet = 3751 psf = 260 psi
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 53
(q minus σ ) (252 psi minus 260 psi) t vo
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 260 psi minus 0 psi = 260 psi
(qt minus σvo)σatm (250 psi minus 260 psi )145 psi = = Qtn prime (σvoσatm)n (260 psi145 psi)n
38
prime σvo 260 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(53)]2
Qtn = 87 n = 10 le 10 Ic = 32 therefore clay
Layer 3
From Figure 10 fs = 20 psi and qc = 4000 psi
qt = qc + u2 ∙ (1 minus a) = 4000 psi + 8 psi ∙ (1 minus 08) = 40016 psi
σvo = 3751 psf + γt ∙ 2 feet = 3751 psf + 1219 pcf ∙ 2 feet = 3995 psf = 277 psi
fs 16 psi Fr() = 100 ∙ = 100 ∙ = 054
(qt minus σvo) (30016 psi minus 277 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 277 psi minus 0 psi = 277 psi
(qt minus σvo)σatm (30016 minus 277)145 psi = = Qtn prime (σvoσatm)n (277 psi145 psi)n
39
primeσvo 277 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(054)]2
Qtn = 1962 n = 052 le 10 Ic = 15 therefore sand
Layer 4
From Figure 10 fs = 2 psi and qc = 250 psi
qt = qc + u2 ∙ (1 minus a) = 250 psi + 0 psi ∙ (1 minus 08) = 250 psi
= 3994 psf + γt ∙ 8 feet = 3994 psf + 1004 pcf ∙ 8 feet = 4798 psf = 333 psi σvo
fs 2 psi Fr() = 100 ∙ = 100 ∙ = 092
(qt minus σvo) (250 psi minus 333 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 333 psi minus 0 psi = 333 psi
(qt minus σvo)σatm (250 psi minus 333 psi )145 psi = = Qtn prime (σvoσatm)n (333 psi145 psi)n
40
prime σvo 333 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(092)]2
Qtn = 65 n = 10 lt 10 Ic = 29 therefore clayey silt
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic Qtn and Fr the first layer is defined as a ldquoDrained Sandrdquo from Figure 17
41
Figure 17 Soil layer 1 using SBT method
Layer 2
Based on values of Ic Qtn and Fr the second layer is defined as a ldquoUndrained Clayrdquo from Figure
18
42
Figure 18 Soil layer 2 using SBT method
43
Layer 3
Based on values of Ic Qtn and Fr the third layer is defined as a ldquoDrained Sandrdquo from Figure 19
Figure 19 Soil layer 3 using SBT method
44
Layer 4
Based on values of Ic Qtn and Fr the fourth layer is defined as a ldquoUndrained Silty Mixrdquo from
Figure 20
Figure 20 Soil layer 4 using SBT method
Effective Stress Friction Angle
Layer 1 (sand)
45
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(4126) = 464deg
Layer 2 (clay)
qt minus σvo
ϕprime(deg) = 295deg ∙ Bq0121 ∙ [0256 + 0336 ∙ Bq + log ( )]
primeσvo
Where
(u2 minus uo) (10 psi minus 0 psi) Bq = = = 004
(qt minus σvo) (252 psi minus 260 psi)
252 psi minus 260 psi ϕprime(deg) = 295deg ∙ 00440121 ∙ [0256 + 0336 ∙ 0044 + log ( )]
260 psi
ϕprime(deg) = 245deg
Layer 3 (sand)
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(1962) = 428deg
Layer 4 (clay)
qt minus σvo
ϕprime(deg) = 295deg ∙ Bq0121 ∙ [0256 + 0336 ∙ Bq + log ( )]
primeσvo
Where
(u2 minus uo) (5 psi minus 0 psi) Bq = = = 002
(qt minus σvo) (251 psi minus 333 psi)
252 psi minus 333 psi ϕprime(deg) = 295deg ∙ 0020121 ∙ [0256 + 0336 ∙ 002 + log ( )]
333 psi
ϕprime(deg) = 265deg
Stress History
46
Layer 1
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (12frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(3000 psi minus 33)072 = 1051 psi
prime σp 1051 psi YSR = = = 642
primeσvo 16 psi
Layer 2
028 028
prime m = 1 minus = 1 minus = 099 25 25 1 + (
Icfrasl ) 1 + (32frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(252 psi minus 260)099 = 734 psi
prime σp 734 psi YSR = = = 28
primeσvo 260 psi
Layer 3
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + ( frasl ) 1 + (15frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(4002 psi minus 277)072 = 1288 psi
47
prime σp 1288 psi YSR = = = 46
primeσvo 277 psi
Layer 4
028 028
prime m = 1 minus = 1 minus = 097 25 25 Ic 1 + ( frasl ) 1 + (29frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(250 psi minus 333)097 = 626 psi
prime σp 626 psi YSR = = = 19
primeσvo 333 psi
Lateral Stress Coefficient
Layer 1
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(464deg)) ∙ 642sin( 464deg) = 56
Layer 2
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(245deg)) ∙ 28sin( 245deg) = 090
Layer 3
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(428deg)) ∙ 46sin( 428deg) = 091
Layer 4
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(266deg)) ∙ 19sin( 266deg) = 073
48
Ground Stiffness and Soil Moduli
Layer 1
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3000 psi minus 16 psi) = 14991 psi
Dprime 14991 psi Eprime = = = 13629 psi
11 11
Eprime 13629 psi Kprime = = = 7571 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(207 MPa)053 + 1355(009 MPa)14 + 236)244 = 2816 MPa = 40873 psi
fs
03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 896 kPa m Vs (mfrasls) = [101 ∙ log(20685 kPa) minus 114]167 ∙ (100 ∙ ) = 2564
20685 kPa s
Vs = 8412 fts
γt 1179 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
2 ft 2 Gmax = ρt ∙ Vs = 37 ∙ (8412 ) = 260 ∙ 106psf = 17992 psi
s
Layer 2
49
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (252 psi minus 260 psi) = 11298 psi
Dprime 11298 psi Eprime = = = 10271 psi
11 11
Eprime 10271 psi Kprime = = = 17118 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(049))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(17 MPa)053 + 1355(008 MPa)14 + 236)244 = 443 MPa = 64309 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 827 kPa m Vs (mfrasls) = [101 ∙ log(17375 kPa) minus 114]167 ∙ (100 ∙ ) = 2646
17375 kPa s
Vs = 8680 fts
γt 1172 pcf slug ρt = = = 36
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 36 ∙ (8680 ) = 274 ∙ 106psf = 19036 psi s
Layer 3
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (4002 psi minus 277 psi) = 19869 psi
Dprime 19869 psi Eprime = = = 18063 psi
11 11
50
Eprime 18063 psi Kprime = = = 10035 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(276 MPa)053 + 1355(014 MPa)14 + 236)244 = 4017 MPa = 58309 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1379 kPa m Vs (mfrasls) = [101 ∙ log(27591 kPa) minus 1379]167 ∙ (100 ∙ ) = 2854
27591 kPa s
Vs = 9363 fts
γt 121 pcf slug ρt = = = 38
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 38 ∙ (9363 ) = 33 ∙ 106psf = 23050 psi s
Layer 4
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (250 psi minus 333 psi) = 1088 psi
Dprime 1088 psi Eprime = = = 989 psi
11 11
Eprime 989 psi Kprime = = = 16490 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(049))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
51
MR = (146(17 MPa)053 + 1355(001 MPa)14 + 236)244 = 360 MPa = 52189 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 138 kPa m Vs (mfrasls) = [101 ∙ log(17238 kPa) minus 114]167 ∙ (100 ∙ ) = 1545
17238 kPa s
Vs = 5069 fts
γt 1004 pcf slug ρt = = = 31
ga 322 ftfrasls2 ft3
2 ft 2 Gmax = ρt ∙ Vs = 31 ∙ (5069 ) = 80 ∙ 105psf = 5565 psi
s
Undrained Shear Strength
Layer 2
qt minus σv0 250 psi minus 26 psi su = = = 187 psi
12 12
Layer 4
qt minus σv0 250 psi minus 333 psi su = = = 181 psi
12 12
Coefficient of Consolidation
52
Example dissipation data are shown in Figure 21 Example calculations are provided
Figure 21 Dissipation t50 data
Layer 1
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (7197)075
cv = = = 359 cmsec 1 sec t50
Layer 2
53
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (94392)075
cv = = = 0008 cmsec 3000 sec t50
Layer 3
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (8303)075
cv = = = 665 cmsec 06 sec t50
Layer 4
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (26714)075
cv = = = 087 cmsec 11 sec t50
Hydraulic Conductivity
Layer 1
cm 1 psi 359 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 10 ∙ 10minus4 cmsec Dprime 14984 psi
Layer 2
cm 1 psi 0008 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 73 ∙ 10minus7 cmsec Dprime 11379 psi
Layer 3
cm 1 psi 665 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 20 ∙ 10minus5 cmsec Dprime 148650 psi
54
Layer 4
cm 1 psi 087 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 18 ∙ 10minus4 cmsec Dprime 10861 psi
55
CHAPTER 3 DIRECT CPT METHOD FOR SHALLOW
FOUNDATIONS
31 PROCEDURE
Shallow foundation analysis is typically done in a two-part traditional procedure The traditional
techniques are no longer required as a direct CPT method for square rectangular and circular shallow
footings is available (Figure 22) This process has the soil types grouped into four main categories sands
silts fissured clays and intact clays When determining soil types for each design it is believed footings
on sands and silts act in a fully drained manner while intact clays act in an undrained manner under
conditions of constant volume In order to determine the vertical stress-displacement-capacity of
square rectangular and circular shallow footings follow the steps provided towards the solution given
by Equation 27
Figure 22 Direct CPT method for shallow foundations
Equation 27 may be used to calculate all footing stresses from zero to the bearing capacity (qmax) To
calculate qmax for a sized footing of width (B) length (L) and thickness (t) follow steps 1 through 7
provided The settlement (s) can be determined after the calculation of qmax by simply rearranging
56
Equation 27 where the allowable stress (qallow) is defined as qmax divided by the factor of safety (FS) For
shallow footings a FS value of 3 is commonly used in geotechnical engineering
120782120787 minus120782120785120786120787 119956 119923 119954119950119938119961 = 119945119956 ∙ 119954119957119951119942119957 ∙ (119913
) ∙ (119913
) 27 119950119938119961
2 119902119898119886119909 0345 1 frasl119865119878 119871 119904 = 119861 ∙ [ ∙ ∙ ( ) ] 28
119861 ℎ119904 119902119905119899119890119905
Where hs = the foundation soil formation parameter qtnet = the net corrected cone tip resistance
311 Step 1 Estimating Footing Dimensions
In Minnesota frost heave can have devastating effect on a shallow foundations It is common practice to
place a foundation bearing elevation below the expected maximum frost depth (roughly 45 to 6 feet
below ground level) With this assumption estimate a footing size (B x L) for design to obtain
representative data from CPT roughly 15middotB below the foundation depth (Df) as shown in Figure 23 The
CPT data collected will consist of
cone tip resistance (qc)
measured porewater pressure acting behind the cone tip (u2) as shown in Figure 24
sleeve friction (fs)
57
Figure 23 Direct CPT method introduction
Figure 24 Differentiation of porewater pressure measurement locations (Lunne et al 1997)
312 Step 2 Soil Characterization
The soil behavior type (SBT) and CPT material index (Ic) govern the value of the formation factor hs The
first step is following the steps in Soil Unit Weight to determine soil layering and γt values for each
layer To determine a representative unit weight (γsoil) for all the layers in the range of Df to 15∙B below
58
Df the user will need to use their engineering judgement on the definition of ldquorepresentative unit
weightrdquo This single value of unit weight is used in further calculations such as the total vertical soil stress (σvo) from Equation 29 and effective vertical stress (σvo) from Equation 30 Values of σvo and σvo
will need to be calculated at 15middotB below Df
120648119959119952 = sum(120632119956119952119946119949 ∙ 119963) 29
prime 120648119959119952 = 120648119959119952 minus 119958119952 30
Where z = Df +15middotB uo=γwatermiddot(z-zw)
The qc will need to be corrected using Equation 31 in the case of fine-grained soils that develop excess porewater pressure during cone penetration These values will be used in the following calculations
119954119957 = 119954119940 + 119958120784 ∙ (120783 minus 119938) 31
Where 119860119899 a = cone area ratio = eg MnDOT commonly uses a = 08 119860119888
119860119899 = cross-sectional area of load cell or shaft 119860119888 = projected area of the cone
The cone area ratio is determined based on the type of piezocone tip used during in-situ field testing
Manufacturer specifications should provide the measured net area ratio (a) for the particular cone
penetrometer as determined by calibration in a pressurized triaxial chamber
313 Step 2a Foundation soil formation parameter
With the representative value γsoil continue to follow the steps in CPT Material Index through Soil
Behavior Type (SBT) to better define the type of soil at depth 15∙B below Df Use the value of Ic
calculated at depth 15∙B below Df to calculate hs with Equation 36 This parameter is based on the soil
type with typical values shown in Figure 25 The data on silts and sands are considered fully drained
whereas the fissured clay subset may be partially drained to undrained
59
23 ℎ119904 = 28 minus
119868119888 15 32
1+( ) 24
Figure 25 Foundation soil formation parameter hs versus CPT material index Ic (Mayne 2017)
314 Step 3 Soil elastic modulus and Poissonrsquos ratio
Details about the soil such as its elastic modulus (Es) and Poissonrsquos ratio (ν) will be needed for further
calculations A representative value for the Es can be determined from in-situ field tests Values of ν can
be assigned as 02 for drained sands and as 05 for undrained loading cases involving clays (Jardine et al
1985 Burland 1989)
315 Step 4 Net cone tip resistance
Calculate qtnet the mean value of net cone tip resistance 15middotB below the foundation bearing elevation
using Equation 33
33 119902119905119899119890119905 = 119902119905 minus 120590119907119900
60
316 Step 5 Bearing capacity of the soil
Use the assumed B and L calculated hs and qtnet to determine the soils bearing capacity (qmax) from
Equation 34 Use Table 6 to calculate qmax by using the maximum allowable settlement ratio (sB)max
correlating to the soil type If hs is in between the given values interpolate to acquire (sB)max
Table 6 Bearing capacity defined by soil type
Type of Soil hs (sB)max
Clean Sands 058 12
Silts 112 10
Fissured Clays 147 7
Intact Clays 270 4
05 minus0345 119904 119871 119902119898119886119909 = ℎ119904 ∙ 119902119905119899119890119905 ∙ (119861
) ∙ (119861
) 34 119898119886119909
317 Step 6 Settlement
Settlement can be calculated directly using the results from Equation 35 A FS equal to 3 is common in
foundation engineering
2 119902119898119886119909 0345 1 frasl119865119878 119871 119904 = 119861 ∙ [ ∙ ∙ ( ) ] 35
119861 ℎ119904 119902119905119899119890119905
318 Step 7 Final Check
Check that the applied stress (q) is less than qmax using Equation 36 Repeat the process again with a
new B andor new L if q gt qmax
05 119904 ) 119902 = ℎ119904 ∙ 119902119905119899119890119905 ∙ (119861
minus0345 119871 ∙ ( )
11986136
61
32 EXAMPLE PROBLEMS
321 Example 3 Direct CPT Method on Sands
A footing size needs to be determined based on the given CPT data collected for the South Abutment
(Figure 27) The footing stress (q) was determined to be 8000 psf Estimate a footing size (B x L) and
determine the bearing capacity of the foundation (qmax) using the direct CPT method provided Also
determine the expected settlement based on the calculated bearing capacity
Figure 26 Diagram of footing profiles for Example 3
The soil elastic modulus determined from seismic CPT (SCPT) are shown in Table 7
Table 7 SCPT Results
Depth (feet) Es (tsf)
Bottom of layer
3 557
6 433
9 557
12 695
17 590
22 501
27 571
32 505
62
Figure 27 CPT data from Northern Minnesota
Solution
Step 1 Assume the footing is placed below the frost depth of 6 feet
Estimate L L = 50 feet = 600 inches
63
Estimate B B = 12 feet = 144 inches
Estimate footing thickness t t = 2 feet
Df = 6 feet = 72 inches
Df + 15middotB = 24 feet = 288 inches
Step 2 Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 28 Schematic of foundation design associated with CPT data
Layer 1
From Figure 28 fs = 20 psi taken as a representative value of the layer
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
64
Layer 2
From Figure 28 fs = 20 psi taken as a representative value of the layer
20psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
Layer 3
From Figure 28 fs = 8 psi taken as a representative value of the layer
8 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1134 pcf
145 psi
Using the unit weights calculated between Df and 15B determine a representative unit weight of the soil to calculate the total and effective soil stresses Based on layer 3 being the weakest supporting layer γsoil = 113 pcf
σvo = γsoil ∙ (119863119891 + 15 ∙ B) = 113 lbsfraslft3 ∙ 24 feet = 2721 psf = 189 psi
prime σvo = σvo minus uo = 2721 psf minus 0 psf = 2721 psf = 189 psi
Calculate the cone tip resistance between Df and Df + 15B below the foundation depth
qt = qc + u2 ∙ (1 minus a) = 1250 psi + 0 psi ∙ (1 minus 08) = 1000 psi
Step 2a CPT Material Index
Using Figure 28 the sleeve friction at 15B below the foundation depth is about 16 psi
65
fs 16 psi Fr() = 100 ∙ = 100 ∙ = 13
(qt minus σvo) (1250 psi minus 189 psi)
Iterate to solve for Ic Steps are not shown for brevity
(qt minus σvo)σatm (1250 psi minus 189 psi )145 psi = = Qtn prime (σvoσatm)n (189 psi145 psi)n
prime σvo 189 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Qtn = 703 n = 072 lt 10 Ic = 210
66
Figure 29 Soil type for example problem 3
Use the Ic value to determine the foundation soil formation parameter
Calculate the foundation soil formation parameter The tip resistance is about 1250 psi at 24 feet and
the cone area ratio was determined to be 08 from information provided by the manufacturer
23 23 hs = 28 minus = 28 minus = 078 = SandSilt 15 15 Ic 210
1 + ( ) 1 + ( ) 24 24
67
Step 3 Determine representative values of soil elastic modulus from soil testing and Poissons ratio The
soil elastic modulus was taken as the average value between depths of Df and 15middotB (6 feet and 24 feet)
433 + 557 + 695 + 590 + 501 Es = = 555 tsf = 708 psi
5
ldquoDrained sandsiltrdquo gives a ν = 020
Step 4 Calculate the net cone tip resistance
qtnet = qt minus σvo = 1250 psi minus 189 psi = 12311 psi
Step 5 Calculate the bearing capacity of the sand
In drained sandssilts the bearing capacity is taken as the stress when (sB) = 011 (or 11 foundation width)
minus0345 minus0345 05 s L 600 qmax = hs ∙ qtnet ∙ ( ) ∙ ( ) = 078 ∙ (9795) ∙ (011)05 ∙ ( )
B B 144
= 193 psi = 27860 psf qmax
Assuming a factor of safety (FS) of 3
qmax = 645 psi = 9287 psf
3
Step 6 Calculate settlement
68
119902119898119886119909 0345 2
1 frasl119865119878 L 119904 = 119861 ∙ [ ∙ ∙ ( ) ]
ℎ119904 119902119905119899119890119905 B
2 0345 1 645 psi 600 inches s = 144 inches ∙ [ ∙ ∙ ( ) ] = 18 inches
078 12311 psi 144 inches
Step 7 Determine if q gt qmax
q = 8000 psf lt 9287psf = qmax3
69
CHAPTER 4 DIRECT CPT METHOD FOR DEEP FOUNDATIONS
41 INTRODUCTION
The axial compression capacity (Qtotal) for a single pile includes a side component (Qside) end bearing
component (Qbase) and pile weight (Wpile) as shown in Figure 30 Since piles commonly push through
several layers a summation of the unit side frictions acting on the pile segments must be considered
over the length of the pile While the examples shown in this Guide resemble hand calculations
computer software is more efficient Programming the procedure is possible and represents a practical
method of designing deep foundations However commercial software is frequently available and is
MnDOTrsquos most common method of designing deep foundations
There are upwards of 40 different direct CPT methods that have been developed over the past five
decades to determine a piles axial compression capacity Many of the earliest methods relied on hand-
recorded information where qc data from mechanical CPTs would be collected at 20 cm intervals
whereas the direct CPT method uses scaled penetrometer readings via specified algorithms to obtain
the pile unit side friction and unit end bearing The method that will be used for deep foundation design
is the Modified UniCone method which uses all three readings of the electronic piezocone (qt fs and u2)
while addressing a variety of pile foundation types
Figure 30 Direct CPT evaluation of axial pile capacity
70
The modified UniCone Method is based upon a total of 330 pile load tests (three times the original
Unicone database) that were associated with SCPTu data Originally the UniCone method provided
approximate soil classification in five groups via a chart of qE vs fs (Figure 31) where qE = qt -u2 Later
using the modified approach with a larger data set provided soil sub classifications as shown in Figure
32 This new 9-zone normalized soil behavior type is determined using CPT data in combination with the
CPT Material Index
Figure 31 UniCone Method soil behavior type using CPT (Mayne 2017)
Figure 32 Modified UniCone Method soil behavior type using CPT (Mayne 2017)
71
42 MODIFIED UNICONE METHOD
In order to determine the axial pile capacity using the modified method the first step requires the
determination of geoparameters as shown in Direct CPT Method for Soil Characterization Once the
soil unit weight and CPT Material index are determined for each soil layer then the effective cone
resistance pile unit side friction (fp) and pile end bearing resistance (qb) can be determined
421 Step 1
Work through the steps provided in CPT Method for Soil Characterization until each soil layer is
defined by its CPT material index and soil behavior type using Figure 6 After these steps have been
completed continue to Step 2 to determine qE qb and fp
422 Step 2
Once qt is determined for each layer the effective cone resistance can be calculated using Equation 37
119902119864 = 119902119905 minus 1199062 37
Where (a) qE is the specific value at each elevation along the pile sides for determining fp and (b) at the bottom of the pile qE is averaged in the vicinity of the pile tip from the tip bearing elevation to about one diameter beneath the tip for determining qb
Using CPT material index and qE the pile end bearing resistance is calculated using Equation 38
38 119902119887 = 119902119864 ∙ 10(0325∙119868119888minus1218)
The pile unit side friction is obtained from qE and Ic
39 119891119901 = 119902119864 ∙ 120579119875119879 ∙ 120579119879119862 ∙ 120579119877119860119879119864 ∙ 10(0732∙119868119888minus3605)
Where θPT = coefficient for pile type (084 for bored 102 for jacked 113 for driven piles) θTC = coefficient for loading direction (111 for compression and 085 for tension) θRATE = rate coefficient applied to soils in SBT zone 1 through 7 (109 for constant rate of penetration test and 097 for maintained load tests)
72
423 Step 3
Due to piles extending through multiple layers the unit side components acting on various pile
segments would need to be summed if not using the direct CPT method Since CPT calculates data at
regular intervals of 2 cm to 5 cm along the sides of the pile the average fp in each layer can be used
directly in Equation 40 to obtain the shaft capacity
119876119904119894119889119890 = 119891119901 ∙ 119860119904 40
Where fp = average pile side friction along pile length from eqn 39 As = πmiddotdmiddotH d = pile diameter and H = length embedded below grade
The base capacity for a pile in compression loading is given by Equation 41 For piles in tension (or uplift)
Qbase can be taken as 0
= 119902119887 ∙ 119860119887 41 119876119887119886119904119890 Where
qb = end bearing resistance from eqn 38 Ab = πmiddotd24 (area of a circular pile)
424 Step 4
The final step is to calculate the axial pile capacity using Equation 42
119876119905119900119905119886119897 = 119876119904119894119889119890 + 119876119887119886119904119890 minus 119882119901119894119897119890 42
43 AXIAL PILE DISPLACEMENTS
Movement of pile foundations can be assessed using elastic continuum theory which has been
developed using finite element analyses boundary elements and analytical closed-form solutions In
the case of piles passing through several soil layers the elastic solution can be used by stacking pile
segments (each with its own stiffness) as represented by soil Youngrsquos modulus The use of software is
recommended for pile groups Several available programs such as DEFPIG GROUP and PIGLET can
handle pile groups under axial and lateralmoment loading
73
44 EXAMPLE PROBLEMS
441 Example 5 Direct CPT Methods Axial Pile Capacity
Several piles need to be placed beneath the edge of a building Use the CPT data collected for this site
shown in Figure 34 to determine the axial capacity for one of the piles Assume round steel driven piles
will be used with a diameter of 1275 inches wall thickness of 025 inches and lengths of 80 feet
Concrete will be used to fill the piles To solve for all geoparameters use the Direct CPT Method for Soil
Characterization section All sand layers can be assumed ldquodrainedrdquo with ν = 02
Figure 33 Deep foundation end bearing pile diagram
74
Figure 34 CPT data from Minnesota for example problem 5
75
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 35 Soil layers using ldquorules of thumbrdquo for pile capacity example using direct CPT method
Layer 1
From Figure 35 fs = 18 psi taken as a representative value of the layer
76
18 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1210 pcf
145 psi
Layer 2
From Figure 35 fs = 12 psi taken as a representative value of the layer
12psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 3
From Figure 35 fs = 20 psi taken as a representative value of the layer
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
CPT Material Index
Layer 1
From Figure 35 fs = 18 psi and qc = 3000 psi
qt = qc + u2 ∙ (1 minus a) = 3000 psi + 20 psi ∙ (1 minus 08) = 3004 psi
∙ 4 feet = 1219 pcf ∙ 4 feet = 4877 psf = 34 psi σvo = γt
fs 18 psi Fr() = 100 ∙ = 100 ∙ = 060
(qt minus σvo) (3004 psi minus 34 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
77
uo = 0 psi
prime σvo = σvo minus uo = 34 psi minus 0 psi = 34 psi
(qt minus σvo)σatm (3004 psi minus 34 psi )145 psi = = Qtn prime )n (σvoσatm (34 psi145 psi)n
prime σvo 34 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(060)]2
Qtn = 3593 n = 038 lt 10 Ic = 14 (ie sand)
Layer 2
From Figure 35 fs = 12 psi and qc = 500 psi
qt = qc + u2 ∙ (1 minus a) = 500 psi + 40 psi ∙ (1 minus 08) = 508 psi
= 4838 psf + γt ∙ 45 feet = 4838 psf + 1172 pcf ∙ 45 feet = 5756 psf = 400 psi σvo
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 260
(qt minus σvo) (508 psi minus 400 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
78
uo = 0 psi
prime σvo = σvo minus uo = 400 psi minus 0 psi = 400 psi
(qt minus σvo)σatm (508 psi minus 400 psi )145 psi = = Qtn prime (σvoσatm)n (400 psi145 psi)n
prime σvo 400 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(260)]2
Qtn = 117 n = 10 le 10 Ic = 29 (ie clayey silt)
Layer 3
From Figure 35 fs = 20 psi and qc = 5000 psi
qt = qc + u2 ∙ (1 minus a) = 5000 psi + 0 psi ∙ (1 minus 08) = 5000 psi
= 5756 psf + γt ∙ 6 feet = 5756 psf + 1219 pcf ∙ 6 feet = 6488 psf = 451 psi σvo
fs 20 psi Fr() = 100 ∙ = 100 ∙ = 040
(qt minus σvo) (5000 psi minus 451 psi)
79
Step 2 Iterate to solve for Ic Steps are not shown for brevity
prime σvo = σvo minus uo = 451 psi minus 0 psi = 451 psi
(qt minus σvo)σatm (5000 psi minus 451 psi )145 psi = = Qtn prime )n (σvoσatm (451 psi145 psi)n
prime σvo 451 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(040)]2
= 1801 n = 06 lt 10 = 15 (ie sand) Qtn Ic
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic (14) Qtn (359) and Fr (06) the first layer is defined as a ldquoGravelly Sandrdquo
from Figure 36
80
Figure 36 Soil layer 1 using SBT method
81
Layer 2
Based on values of Ic (29) Qtn (117) and Fr (26) the second layer is defined as a ldquoSilty Mixrdquo
from Figure 37
Figure 37 Soil layer 2 using SBT method
82
Layer 3
Based on values of Ic (15) Qtn (180) and Fr (04) the third layer is defined as a ldquoSandrdquo from
Figure 38
Figure 38 Soil layer 3 using SBT method
Effective Cone Resistance
Layer 1
83
qE = qt minus u2 = 3004 psi minus 20 psi = 2984 psi
Layer 2
qE = qt minus u2 = 508 psi minus 40 psi = 468 psi
Layer 3
qE = qt minus u2 = 5000 psi minus 0 psi = 5000 psi
Pile End Bearing Resistance
Layer 3
= qE ∙ 10(0325∙Icminus1218) qb = 5000 psi ∙ 10(0325∙15minus1218) = 9085 psi
Pile Unit Side Friction
Layer 1
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 2984 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙14minus3605) = 99 psi
Layer 2
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 468 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙29minus3605) = 211 psi
Layer 3
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 5000 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙15minus3605) = 202 psi
84
Axial Pile Capacity
The side capacity is based on pile design as shown in Figure 39
Layer 1
Qside = fp ∙ As = 99 psi ∙ (π ∙ d ∙ L) = 99 psi ∙ (π ∙ 1275 in ∙ 48 in) = 19064 lb
Layer 2
Qside = fp ∙ As = 211 psi ∙ (π ∙ d ∙ L) = 211 psi ∙ (π ∙ 1275 in ∙ 588 in) = 457122 lb
Layer 3
Qside = fp ∙ As = 202 psi ∙ (π ∙ d ∙ L) = 202 psi ∙ (π ∙ 1275 in ∙ 240 in) = 58170 lb
Figure 39 Soil layering compared to pilings
85
End Bearing Resistance in Layer 3
d2 1275 in2
Qbase = qb ∙ Ab = 9085 psi ∙ (π ∙ ) = 9085 psi ∙ (π ∙ ) = 115932 lb 4 4
Wp = (γsteel ∙ Asteel ∙ Depth) + (γsteel ∙ Aconc ∙ Depth)
Wp = (490 pcf ∙ 003 ft2 ∙ 55 ft) + (150 pcf ∙ 085 ft2 ∙ 55 ft) = 7724 lbs
Layer 3
Qtotal = ΣQside + Qblayer3 minus Wp
Qtotal = (19064 + 45713 + 58170) lbs + 115932 lbs minus 7724 lbs = 642563 lbs
= 642 kips Qtotal
Table 8 Summary of selected and calculated parameters
Layer L (in) qc
(psi)
fs
(psi)
Fr Ic Qtn qE
(psi)
qb
(psi)
Qbase
(lb)
fp
(psi)
Qside
(lb)
Qtotal
(kip)
1 48 3000 18 06 14 356 2984 NA NA 99 19064
2 588 500 12 26 29 117 468 NA NA 211 457122
3 240 5000 20 04 15 180 5000 9085 115932 202 58170
Total 115932 534355 642
86
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Agaiby SA (2018) Advancements in the interpretation of seismic piezocone tests in clays and other geomaterials PhD dissertation School of Civil amp Environmental Engineering Georgia Institute of Technology Atlanta GA
Agaiby SS and Mayne PW (2018) Interpretation of piezocone penetration and dissipation tests in sensitive Leda Clay at Gloucester Test Site Canadian Geotechnical Journal (accepted for publication 08 March 2018) CGJ-2017-0388
Amar S Baguelin F Canepa Y and Frank R (1998) New design rules for the bearing capacity of shallow foundations based on Menard PMT Geotechnical Site Characterization Vol 2 (Proc ISC-1 Atlanta) 727-733
American Petroleum Institute (API) (1981) Planning designing and constructing fixed offshore platforms Recommended practice API-RP2A 17th edition Washington DC API
American Petroleum Institute (API) (1987) Planning designing and constructing fixed offshore platforms Recommended practice API-RP2A 18th edition Washington DC API
American Petroleum Institute (API) (1989) Planning designing and constructing fixed offshore platforms Recommended practice API-RP2A 19th edition Washington DC API
Andersen KH and Stenhamar P (1982) Static plate loading tests on overconsolidated clay Journal of the Geotechnical Engineering Division ASCE 108 (GT7) 919-934
Basu D and Salgado R (2012) Load and resistance factor design of drilled shafts in sands Journal of Geotechnical amp Geoenvironmental Engineering 138 (12) 1455-1469
Berardi R Jamiolkowski M and Lancellotta R (1991) Settlement of shallow foundations in sands selection of stiffness on the basis of penetration resistance Geotechnical Engineering Congress Vol 1 185-200 (Proc GSP-27 Boulder) Reston Virginia ASCE
Brand EW Muktabhant C and Taechathummarak A (1972) Load tests on small foundations in soft clay Performance of Earth and Earth-Supported Structures Vol 1 Part 2 903-928 (Proc PC) ASCE Reston Virginia ASCE
Briaud JL and Gibbens RM (1999) Behavior of five large spread footings in sand Journal of Geotechnical amp Geoenvironmental Engineering 125 (9) 787-797
Briaud JL (2007) Spread footings in sand load-settlement curve approach Journal of Geotechnical amp Geoenvironmental Engineering 133 (8) 905-920
Brown DA Turner JP and Castelli RJ (2010) Drilled Shafts Construction Procedures and LRFD Design Methods Geotechnical Engineering Circular GEC 10 Report No FHWA NHI-10-016 Washington DC National Highway Institute US DOT
87
Burland J B (1973) Shaft Friction of Piles in Clay -A Simple Fundamental Approach Ground Eng 6(3) 30-42
Cai G Liu S and Puppala A (2012) Reliability assessment of CPTu-based pile capacity predictions in soft clay deposits Engineering Geology 84-91
Canadian Geotechnical Society (1992) Canadian Foundation Engineering Manual Third Edition Vancouver British Columbia BiTech Publishers
Cerato AB and Lutenegger AJ (2007) Scale effects of shallow foundation bearing capacity on granular material Journal of Geotechnical amp Geoenvironmental Engineering 133 (10) 1192-1201
Chen BS-Y and Mayne PW (1996) Statistical relationships between piezocone measurements and stress history of clays Canadian Geotechnical Journal 33 (3) 488-498
Cho W and Finno RJ (2010_ Stress-strain responses of block samples of compressible Chicago glacial clays Journal of Geotechnical amp Geoenvironmental Engineering 136 (1) 178-188
Chung SF Randolph MF and Schneider JA (2006) Effect of penetration rate on penetrometer resistance in clay Journal of Geotechnical amp Geoenvironmental Engineering 132 (9) 1188-1196
Clausen CJF Aas PM and Karlsrud K (2005) Bearing capacity of driven piles in sand the NGI approach Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis Group
Consoli NC Schnaid F and Milititsky J (1998) Interpretation of plate load tests on residual soil site Journal of Geotecnical amp Geoenvironmental Engingeering 124 (9) 857-867
Dasenbrock DD (2006) Assessment of pile capacity by static and dynamic methods to reconcile design predictions with observed performance Proc 54th Minnesota Geotechnical Engineering Conference 95-108 edited by Labuz J Univ of Minnesota St Paul
DeBeer E and Martens A (1957) Method of computation of an upper limit for the influence of heterogeneity of sand layers on the settlement of bridges 275-282 Proc ICSMFE-4 Vol 1 London ICSMFE
DeBeer EE (1965) Bearing capacity and settlement of shallow foundations on sand Proc Symposium on Bearing Capacity and Settlement of Foundations 15-33 Duke University Durham NC
Decourt L (1999) Behavior of foundations under working load conditions Proc PCSMGE-11 Foz do Iguassu Brazil Vol 4 453-487
Demers D and Leroueil S (2002) Evaluation of preconsolidation pressure and the overconsolidation ratio from piezocone tests of clay deposits in Quebec Canadian Geotechnical Journal 39 (1) 174-192
Eslaamizaad S and Robertson PK (1996) Cone penetration test to evaluate bearing capacity of foundations in sand 429-438 Proc CGCFG-49 Vol 1 St Johns Newfoundland
Eslami A Alimirzaei M Aflaki E and Molaabasi H (2017) Deltaic soil behavior classification using CPTu records Marine Georesources amp Geotechnology 35 (1) 62-79
88
Eslami A and Gholami M (2005) Bearing capacity analysis of shallow foundations from CPT data 1463-1466 Proc ICSMGE- 16 Vol 3 (Osaka) Millpress Rotterdam
Eslami A and Gholami M (2006) Analytical model for the ultimate bearing capacity of foundations from cone resistance Scientia Iranica 13 (3) 223-233
Eslami A and Fellenius BH (1997) Pile capacity by direct CPT and CPTu methods applied to 102 case histories Canadian Geotechnical Journal 34 (6) 886-904
Fellenius BH (2009) Basics of Foundation Design Vancouver BiTech Publishers
Fellenius BH and Altaee A (1994) Stress and settlement of footings in sand Vertical and Horizontal Deformations of Foundations and Embankments 2 1760-1773
Fellenius BH and Eslami A (2000) Soil profile interpreted from CPTu data Proc Geotechnical Engineering Conference Year 2000 Geotechnics Asian Inst of Technology Bangkok Thailand Available from wwwfelleniusnet
Finno R J and Chung C K (1992) Stress-strain-strength responses of compressible Chicago glacial clays Journal of Geotechnical Engineering 118 (10) 1607ndash1625
Fleming K Weltman A Randolph M and Elson K (2009) Piling Engineering Third Edition London Taylor amp Francis Group
Frank R and Magnan J-P (1995) Cone penetration testing in France national report Proceedings International Symposium on Cone Penetration Testing 3 (CPT95 Linkoumlping) Swedish Geotechnical Society Report 3(95) 147-156
Gavin K Adekunte A and OKelly B (2009) A field investigation of vertical footing response on sand Geotechnical Engineering 162 257-267
Hegazy YA (1998) Delineating geostratigraphy by cluster analysis of piezocone data PhD dissertation School of Civil amp Environmental Engineering Georgia Institute of Technology Atlanta GA
Holtz RD Kovacs WD and Sheehan TC (2011) An Introduction to Geotechnical Engineering 2nd Edition London Pearson Publishing
Hunt CE Pestana JM Bray JD and Riemer M (2002) Effect of pile driving on static and dynamic properties of soft clay Journal of Geotechnical amp Geoenvironmental Engineering 128 (1) 13-24
Jamiolkowski M Ladd CC Germaine JT and Lancellotta R (1985) New developments in field and lab testing of soils Proc ICSMFE-11 1 57-154
Jardine R Chow F Overy R and Standing J (2005) ICP design methods for driven piles in sands and clays London Thomas Telford Publishing
Jardine RJ Lehane BM Smith P and Gildea P (1995) Vertical loading experiments on rigid pad foundations at Bothkennar Geotechnique 45 (4) 573-597
89
Karlsrud K (2012) Prediction of load-displacement behavior and capacity of axially-loaded piles in clay based on analyses and interpretation of pile load test results PhD Dissertation Norwegian Univ Science amp Technology Trondheim Norway
Karlsrud K Clausen CJF and Aas PM (2005) Bearing capacity of driven piles in clay the NGI approach Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis
Karlsrud K Lunne T Kort DA and Strandvik S (2001) CPTU correlations for clays Proc 16th ICSMGE 2 683-702
Kimura T Kusakabe O and Saitoh K (1985) Geotechnical model tests of bearing capacity problems in a centrifuge Geotechnique 35 (1) 33-45
Knappett JA and Craig RF (2012) Craigs Soil Mechanics 8th Edition London Spon Press Taylor amp Francis
Kolk HJ and Velde EVD (1996) A reliable method to determine friction capacity of piles driven into clays Paper No 7993 Proc Offshore Technology Conference Houston
Kolk HJ Baaijens AE and Senders M (2005) Design criteria for pipe piles in silica sands Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis Group
Kulhawy FH (2004) On the axial behavior of drilled shafts Geosupport 2004 1-18
Kulhawy FH and Jackson CS (1989) Foundation Engineering Current Principles amp Practices 2 (GSP 22) Reston VA ASCE
Kulhawy FH and Mayne PW (1990) Manual for Estimating Soil Properties for Foundation Design EPRI Report EL-6800 Electric Power Research Institute Palo Alto 306 page Download from wwwepricom
Kulhawy FH Trautmann CH Beech JF ORourke TD and McGuire W (1983) Transmission Line Structure Foundations for Uplift-Compression Loading Report EL-2870 Palo Alto CA Electric Power Research Institute wwwepricom
Ku T and Mayne PW (2015) In-situ lateral stress coefficient (K0) from shear wave velocity measurements in soils Journal of Geotechnical amp Geoenvironmental Engineering 141 (12) 101061(ASCE) GT1943-56060001354
Ladd CC and DeGroot DJ (2003) Recommended practice for soft ground site characterization Soil amp Rock America 2003 3-57
Larsson R (1997) Investigations and Load Tests in Silty Soils SGI Report R-54 Linkoumlping Swedish Geotechnical Institute
Larsson R (2001) Investigations and Load Tests in Clay Till SGI Report R-59 Linkoumlping Swedish Geotechnical Institute
Lee J and Salgado R (2005) Estimation of bearing capacity of circular footings on sands based on cone penetration tests Journal of Geotechnical amp Geoenvironmental Engineering 131 (4) 442-452
90
Lehane BM (2003) Vertically loaded shallow foundation on soft clayey silt Geotechnical Engineering 156 (1) Thomas Telford London 17-26
Lehane BM (2013) Relating foundation capacity in sands to CPT qc Geotechnical amp Geophysical Site Characterization 1 London Taylor amp Francis Group
Lehane BM Doherty JP and Schneider JA (2008) Settlement prediction for footings on sand Deformational Characteristics of Geomaterials (IS-Atlanta) 1 Rotterdam Millpress
Lehane BM Li Y and Williams R (2013) Shaft capacity of displacement piles in clay using the CPT Journal of Geotechnical amp Geoenvironmental Engineering 139 (2) 253-266
Lehane BM Schneider JA and Xu X (2005) The UWA-05 method for prediction of axial capacity of driven piles in sand Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis
Low HE Lunne T Andersen KH Sjursen MA Li X and Randolph MF 2010 Estimation of intact and remoulded undrained shear strengths from penetration tests in soft clays Geotechnique 60 (11) 843-859
Lunne T Robertson PK and Powell JJM (1997) Cone Penetration Testing in Geotechnical Practice London RoutledgeBlackie Academic
Lutenegger AJ and DeGroot DJ (1995) Settlement of Shallow Foundations on Granular Soils Report Contract 6332 Task Order 4 submitted to Massachusetts Highway Dept by University of Mass-Amherst Amherst MA
Lutenegger AJ and Adams MT (2003) Characteristic load-settlement behavior of shallow foundations Proc FONDSUP 2 381-392
Mase T and Hashiguchi K (2009) Numerical analysis of footing settlement problem by subloading surface model Soils amp Foundations 49 (2) 207-220
Mayne PW (1987) Determining preconsolidation stress and penetration pore pressures from DMT contact pressures Geotechnical Testing Journal 10(3) 146-150
Mayne PW (1991) Determination of OCR in Clays by Piezocone Tests Using Cavity Expansion and Critical State Concepts Soils and Foundations 31 (2) 65-76
Mayne PW (2007) NCHRP Synthesis 368 on Cone Penetration Test Washington DC Transportation Research Board National Academies Press wwwtrborg
Mayne PW (2009) Geoengineering Design Using the Cone Penetration Test Richmond BC Canada ConeTec Inc
Mayne PW (2014) Keynote Interpretation of geotechnical parameters from seismic piezocone tests Proceedings 3rd Intl Symp on Cone Penetration Testing 47-73
Mayne PW (2016) Evaluating effective stress parameters and undrained shear strengths of soft-firm clays from CPT and DMT Australian Geomechanics Journal 51 (4) 27-55
91
Mayne PW (2017) Stress history of soils from cone penetration tests 34th Manual Rocha Lecture Soils amp Rocks Vol 40 (3) Saouml Paulo ABMS
Mayne PW Christopher B Berg R and DeJong J (2002) Subsurface Investigations -Geotechnical Site Characterization Publication No FHWA-NHI-01-031 Washington DC National Highway Institute Federal Highway Administration
Mayne PW Coop MR Springman S Huang A-B and Zornberg J (2009) Geomaterial Behavior and Testing Proc 17th Intl Conf Soil Mechanics amp Geotechnical Engineering 4 Rotterdam MillpressIOS Press
Mayne PW and Illingworth F (2010) Direct CPT method for footings on sand using a database approach Proc ISCPT-2 3 315-322
Mayne PW and Niazi F (2017) Recent developments and applications in field investigations for deep foundations Proceedings 3rd Bolivian Conference on Deep Foundations 1 141-160
Mayne PW Peuchen J and Baltoukas D (2015) Piezocone evaluation of undrained strength in soft to firm offshore clays Frontiers in Offshore Geotechnics III 2 1091-1096
Mayne PW and Poulos HG (1999) Approximate displacement influence factors for elastic shallow foundation systems ASCE Journal of Geotechnical amp Geoenvironmental Engineering 125 (6) 453-460
Mayne PW Uzielli M and Illingworth F (2012) Shallow footing response on sands using a direct method based on cone penetration tests Full Scale Testing and Foundation Design (GSP 227) Reston Virginia ASCE
Mayne PW and Woeller DJ (2014) Direct CPT method for footings on sands silts and clays Geocharacterization for Modeling and Sustainability (GSP 234) 1983-1997
McClelland B (1974) Design of Deep Penetration Piles for Ocean Structures Journal Geot Eng Div 100(GT7) 705-747
Mesri G (1975) Discussion on new design procedure for stability of soft clays ASCE Journal of the Geotechnical Engineering Division 101(GT4) pp 409-412
Meyerhof GG (1956) Penetration tests and bearing capacity of cohesionless soils Journal of Soil Mechanics amp Foundations Division 82 (SM1) 1-19
Meyerhof GG (1974) General Report Penetration testing outside Europe Proceedings ESOPT-1 Vol 21 Swedish Geotechnical Society Stockholm
Meyerhof GG (1976) Bearing capacity and settlement of pile foundations Journal of Geotechnical Engineering Division 102(3) 197-228
Missiaen T Verhegge J Heirman K and Crobe P (2015) Potential of cone penetrating testing for mapping deeply buried palaeolandscapes in the context of archaeological surveys in polder areas Journal of Archaeological Science 55 174-187
92
Mlynarek Z Wierzbicki J Goglik S and Bogucki M (2014) Shear strength and deformation parameters of peat and gyttja from CPTu DMT and VST Proceedings 5th International Workshop on CPTu and DMT in soft clays and organic soils 193-210 Poznan Poland Polish Committee on Geotechnics Menard Polska Tensar Internationa
Nejaim PF Jannuzzi GMF and Danziger FAB (2016) Soil behavior type of the Sarapuiacute II test site Geotechnical amp Geophysical Site Characterization 5 1009-1014
Niazi FS and Mayne PW (2013) Cone penetration test based direct methods for evaluating static axial capacity of single piles Geotechnical and Geological Engineering 31 (4) 979-1009
Niazi FS and Mayne PW (2015) Enhanced Unicone expressions for axial pile capacity evaluation from piezocone tests Proc International Foundations Conference amp Equipment Expo RestonVA ASCE
Niazi FS and Mayne PW (2016) CPTu-based enhanced UniCone method for pile capacity Engineering Geology 212 21-34
Nowacki F Karlsrud K and Sparrevik P (1996) Comparison of recent tests on OC clay and implications for design Large Scale Pile Tests in Clay London Thomas Telford
ONeill MW (2001) Side resistance in piles and drilled shafts Journal of Geotechnical amp Geoenvironmental Engineering 127 (1) 3-16
Ouyang Z and Mayne PW (2018) Effective friction angle of clays and silts from piezocone Canadian Geotechnical Journal in press doiorg101139cgj-2017-0451
Paikowsky SG Canniff MC Lesny K Kisse A Amatya S and Muganga R (2010) LRFD Design and Construction of Shallow Foundations for Highway Bridge Structures NCHRP Report 651 Washington DC National Cooperative Highway Research Program Transportation Research Board
Pestana JM Hunt CE and Bray JD (2002) Soil deformation and excess pore pressure filed around a closed-ended pile Journal of Geotechnical amp Geoenvironmental Engineering 128 (1) 1-12
Poulos HG and Davis EH (1974) Elastic Solutions for Soil and Rock Mechanics New York Wiley amp Sons
Poulos HG and Davis E (1980) Pile Foundation Analysis amp Design New York John Wiley amp Sons
Powell JJM and Lunne T (2005) Use of CPTU data in clays and fine-grained soils Studia Geotechnica et Mechanica XXVII(3) 29-66
Randolph MF (2003) Rankine Lecture Science and empiricism in pile foundation design Geotechnique 53 (10) 847-875
Robertson PK (1990 1991) Soil classification using the cone penetration test Canadian Geotechnical Journal 27 (1) 151-158 Closure 28 (1) 176-178
Robertson PK (1991) Estimation of foundation settlements in sand from CPT Geotechnical Engineering Congress 2 Reston Virginia ASCE
93
Robertson PK (2009) Cone penetration testing a unified approach Canadian Geotechnical J 46 (11) 1337-1355
Robertson PK (2009a) Performance based earthquake design using the CPT Keynote Lecture International Conference on Performance-Based Design in Earthquake Geotechnical Engineering IS Tokyo Tsukuba Japan
Robertson PK (2009b) Interpretation of cone penetration tests a unified approach Canadian Geotechnical Journal 46 (11) 1337-1355
Robertson PK and Cabal KL (2007) Guide to cone penetration testing Second Edition Signal Hill California Gregg In-Situ
Robertson PK and Cabal KL (2015) Guide to Cone Penetration Testing for Geotechnical Engineering 6th Edition Signal Hill California Gregg Drilling amp Testing Inc
Schmertmann JH (1970) Static cone to compute static settlement over sand Journal of the Soil Mechanics and Foundations Division 95 (SM3) 1011-1043
Schmertmann JH (1978) Guidelines for cone penetration test performance and design Report FHWA-TS-78-209 Washington DC Federal Highway Administration
Schnaid F Wood WR Smith AKC and Jubb P (1993) Investigation of bearing capacity and settlements of soft clay deposits at Shellhaven Predictive Soil Mechanics London Thomas Telford
Schneider JA Xu X and Lehane BM (2008) Database assessment of CPT-based design methods for axial capacity of driven piles in siliceous sands J Geotechnical and Geoenvironmental Engineering 134 (9) 1227-1244
Semple D (1980) Recent Developments in the Design amp Construction of Piles London Institution of Civil Engineers
Senneset K Sandven R amp Janbu N (1989) Evaluation of soil parameters from piezocone tests In-Situ Testing of Soil Properties for Transportation Transportation Research Record 1235 24-37
Shahri AA Melehmir A and Juhlin C (2015) Soil classification analysis based on piezocone penetration test data Engineering Geology 189 32-47
Stuedlein AW and Holtz RD (2010) Undrained displacement behavior of spread footings in clay The Art of Foundation Engineering Practice GSP 198 Reston Virginia ASCE
Takesue K Sasao H and Matsumoto T (1998) Correlation between ultimate pile skin friction and CPT data Geotechnical Site Characterization 2 1177ndash1182
Tand KE Funegard EG and Briaud J-L (1986) Bearing capacity of footings on clay CPT method Use of In-Situ Tests in Geotechnical Engineering GSP 6 1017-1033
Tand KE Warden PE and Funegaringrd EG (1995) Predicted-measured bearing capacity of shallow footings on sand PISCPT 2 589-594
Thomas D (1968) Deep soundings test results and the settlement of spread footings on normally consolidated sands Geotechnique 18 (2) 472-488
94
Tomlinson MJ (1957) The adhesion of piles driven into clay soils Proc 4th Intl Conf on Soil Mechanics and Foundation Engineering 2 66-71
Tomlinson MJ (1986) Foundation Design amp Construction London Longman Scientific
Tomlinson MJ (1987) Pile Design and Construction Practice London Viewpoint Publication
Trofimenkov JG (1974) Penetration testing in the USSR Proceedings ESOPT-1 1 147-154
Tumay MT Hatipkarasulu Y Marx ER and Cotton B (2013) CPTPCPT-based organic material profiling Proc 18th Intl Conf on Soil Mechanics amp Geotechnical Engineering Paris 633-636
Uzielli M and Mayne PW (2011) Statistical characterization and stochastic simulation of load-displacement behavior of shallow footings Proc Georisk 2011 Geotechnical Risk Management amp Assessment (GSP 224) 672-679
Uzielli M and Mayne PW (2012) Load-displacement uncertainty of vertically-loaded shallow footings on sands and effects on probabilistic settlement estimation Georisk Assessment and Management of Risk for Engineered Systems and GeoHazards 6 (1) 50-69
Valsson SM (2016) Detecting quick clay with CPTu Proceedings 17th Nordic Geotechnical Meeting Reykajavik Iceland Challenges in Nordic Geotechnics
Van Dijk BFJ and Kolk HJ (2011) CPT-based design method for axial capacity of offshore piles in clay Frontiers in Offshore Geotechnics II 555-560
Vesić AS (1975) Bearing capacity of shallow foundations Foundation Engineering Handbook New York Van Nostrand Reinhold Publishers
Vesic AS (1977) NCHRP Synthesis of Highway Practice 42 Design of Pile Foundations Washington DC Transportation Research Board
Yu F and Yang J (2012) Base capacity of open-ended steel pipe piles in sand J Geotechnical and Geoenvironmental Engineering 138 (9) 1116-1128
Zawrzykraj P Rydelek P and Bakowska A (2017) Geoengineering properties of Eemian peats from Radsymin (central Poland) in the light of static cone penetration and dilatometer tests Engineering Geology 226 290-300
95
APPENDIX A
List of Figures
Figure A1 Conventional methods versus direct CPT approach to shallow foundation response A-3
Figure A2 Direct bearing capacity relationship for embedded square and strip footings situated on clean
sands (after Schmertmann 1978) A-6
Figure A3 Direct CPT bearing factor Rk as function of footing width B and embedment depth from finite
element analyses (Tand et al 1995) A-6
Figure A4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio and sand
consistency (Eslaamizaad amp Robertson 1996) A-7
Figure A5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp Salgado (2005) in
terms of footing size relative density and base settlement A-8
Figure A6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and normalized
cone tip resistance (after Eslami amp Gholami 2005) A-8
Figure A7 Direct relationship between ultimate bearing stress in clays and measured cone tip resistance
(Schmertmann 1978) A-10
Figure A8 Direct relationship between ultimate foundation bearing stress in clays and net cone tip
resistance (after Tand et al 1986) A-11
Figure A9 Measured load-displacements for each of the five large footings at TAMU (Briaud amp Gibbens
1994) sand site A-12
Figure A10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU (Briaud amp
Gibbens 1994) footings A-13
Figure A11 Characteristic stress versus square root normalized displacements for TAMU footings A-15
Figure A12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sand A-16
Figure A13 Summary of sand formation factors with corresponding CPT resistances A-19
Figure A14 Normalized foundation stresses vs square root of normalized displacement for spread
footings on sand (modified after Mayne et al 2012) A-20
Figure A15 Definitions of capacity from three types of observed foundation load test responses
(modified after Kulhawy 2004) A-21
Figure A16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footings A-22
A-1
Figure A17 Measured load vs displacement curves for 3 shallow foundations on clay at Baytown Texas
(Stuedlein amp Holtz 2010) A-25
Figure A18 Characteristic stress vs square root of normalized displacement for all three foundations at
Baytown site (Stuedlein amp Holtz 2010) A-26
Figure A19 Normalized foundation stress vs square root of normalized displacements A-27
Figure A20 Applied foundation stress vs CPT-calculated stress for 67 footings A-28
Figure A21 Applied footing stress vs CPT calculated stress on logarithmic scales A-29
Figure A22 Summary graph and equations of direct CPT method for evaluating the vertical A-30
Figure A23 Foundation soil formation parameter hs versus CPT material index Ic A-32
Figure A24 Bearing capacity ratio qmaxqnet versus CPT material index Ic A-33
Figure A25 Influence factor for rectangular foundations from elastic theory solution (Mayne amp
Dasenbrock 2017) A-34
List of Tables
Table A1 Direct CPT methods for bearing capacity of footings on clean sands A-4
Table A2 Direct CPT methods for bearing capacity of footings on clays A-9
Table A3 Summary of large footings soil conditions and reference sources of database A-17
Table A4 Sources for shallow foundation performace A-34
A-2
A1 INTERPRETATION OF SOIL PARAMETERS FROM CONE
PENETRATION TESTS
A11 Introduction
Geotechnical site characterization is an important first step towards the evaluation of
subsurface conditions and determination of soil layering geomaterial classification and the
evaluation of soil engineering parameters for the analysis and design of foundations retaining
walls tunnels excavations embankments and slope stability Increasing use of the electronic
cone penetrometer in highway applications is prominent in the USA because the results are
obtained much faster and less expensive than traditional methods that rely on rotary drilling
augering sampling and laboratory testing Moreover the cone penetration test (CPT) provides
at least three independent and continuous readings on soil behavior that are digitally recorded
and fully available at the end of the sounding As such the reliable and consistent
interpretation of CPT data is important since civil engineering works will best realize the
efficiency and economy of this technology in constructed facilities for county state and
interstate projects
A111 Cone Penetration Testing
The cone penetrometer is an electronically-instrumented steel probe that is vertically pushed into the
ground at constant rate of 20 mms (08 insec) using a hydraulic system The penetrometer is outfitted
with load cells pressure transducers and inclinometers that measure at least three readings
continuously with depth (a) cone tip resistance qt (b) sleeve friction fs and (c) porewater pressure u2
With the latter reading the device is called a piezocone thus the designation CPTu is used to indicate
porewater pressure Additional sensors are available that can be used to measure inclination
resistivity pH temperature shear wave velocity and other readings if desired
Figure A1 shows a schematic rendering of the cone penetration test that is performed in the
field per ASTM D 5778 procedures The hydraulic system is nominally rated at 180 kN (20 tons)
capacity and often mounted on a truck but can also be positioned on tracked vehicles or
anchored frames Figure A2 shows a MnDOT cone truck that uses the full dead weight of the
rig for the hydraulic reaction forces which are applied at the center of the vehicle The cab is
enclosed so that soundings may proceed during inclement weather and the data acquisition
system and operator are protected
Figure A 1 Setup and procedures for co ne penetration testing (CPT)
A-4
Figure A2 CPT truck-mounted rig used by MnDOT
A representative sounding taken at the I-35 test site just northeast of Saint Paul is presented in Figure
A3 Here the separate profiles of qt fs and u2 with depth are shown in side-by-side plots In the last
graph the hydrostatic porewater pressure (uo) due to the groundwater table is indicated by the blue
dashed line
These readings are used to interpret the soil layers types and properties of the ground as
discussed in the following sections
A-5
Figure A3 Representative CPT sounding at I-35 test site northeast of St Paul MN showing profiles of
(a) cone resistance (b) sleeve friction and (c) penetration porewater pressures
A112 Soil Parameter Interpretation
A variety of soil engineering parameters can be interpreted from CPT results on the basis of
theoretical frameworks analytical models or numerical simulations otherwise by empirical
methods based on correlations and statistics of databases Figure A4 shows a conceptual
utilization of the readings from the CPT for interpretation of geostratigraphy compressibility
flow characteristics and stress-strain-strength behavior of soils
Table A1 lists a selection of geoparameters that have been addressed for these purposes The
most common or useful relationships will be discussed in subsequent sections and reference is
made to other sources (Lunne et al 1997 Mayne 2007 Robertson amp Cabal 2015) for additional
details
A-6
Figure A4 Conceptual post-processing of CPT results for geoparameter evaluation
Table A1 List of common geoparameters interpreted from CPT data
Symbol Parameter Remarks Notes
SBT Soil behavior type (SBT)
Ic CPT material index
γt Unit weight
ρt Mass Density
σvo Overburden stress
u0 Hydrostatic pressure
A-7
σvo Effective stress
Table A1 Continued
Symbol Parameter Remarks Notes
σp Preconsolidation stress σp = 033 (qnet)m where m = fctn (Ic)
(or Effective Yield Stress) where stresses are in kPa
YSR Yield stress ratio YSR = σpσvo (formerly OCR)
c Effective cohesion intercept
ϕ Effective friction angle
su Undrained shear strength
D Constrained modulus
G0 Small-strain modulus
G Shear modulus
E Youngs modulus
cvh Coefficient of consolidation
K Bulk modulus
A-8
LI
K0 Lateral stress coefficient
k Hydraulic conductivity
Liquefaction index
A12 Geostratigraphy and Soil Behavioral Type (SBT)
In routine CPT soil samples are not normally taken and thus the measured stress friction andor
porewater pressure readings are used to infer the types of soils This can be accomplished using
three basic approaches
a Rules of Thumb or approximate guidelines for a quick visual assessment
b Soil Behavioral Type (SBT) Charts that are based on either the three readings (qt fs u2) net readings
including net cone resistance (qnet = (qt-σvo) effective cone resistance (qE = qt - u2) friction ratio (Rf =
100middotfsqt) and excess porewater pressures or using normalized CPT readings such as Q F Bq or U
c Probabilistic methods where the CPT readings have been calibrated from lab tests on
recovered soil samples (eg Tumay et al 2013)
A121 Approximate Rules of Thumb
The approximate rules of thumb provide simple guidelines for soil type and suggest that sands are
identified when qt gt 5 MPa and u2 asymp u0 whereas intact clays are prevalent when qt lt 5 MPa and u2 gt u0
(Mayne et al 2002) The magnitude of porewater pressures reflects the consistency of the intact clay
such that soft (u2 asymp 2middotu0) firm (u2 asymp 4middotu0) stiff (u2 asymp 8middotu0) and hard (u2 asymp 20middotu0) However for fissured
overconsolidated clays the measured porewater pressures are less than hydrostatic in fact often
negative u2 lt 0 On land negative porewater pressure readings at the shoulder filter element of the
piezocone should be greater than -1 bar ie u2 gt -100 kPa (-147 psi) Also for clean sands the friction
ratio (FR = 100middotfsqt lt 1) while for insensitive clays (FR gt 4)
A-9
Figure A5 presents a 36-m deep piezocone sounding from the MnDOT Wakota Bridge site which crosses
the Mississippi River at I-494 (Dasenbrock 2006) The qt and u2 readings have been annotated using the
aforementioned rules of thumb indicating the predominance of sands at this site As noted there are
five interbedded clay layers evident at depths of 1 m 3-4 m 7 - 12 m 17 m and 22-27 m
Figure A5 Interpretation of soil types from CPT data taken from the Wakota Bridge on I-494 using
approximate rules of thumb
A122 Soil Behavioral Type Charts
The most common approach for identification of soil types is based on soil behavioral charts
Reviews of several chart methods are provided elsewhere (Kulhawy amp Mayne 1990 Fellenius amp
Eslami 2000 Shahri et al 2015) Current popularity favors the 9-zone classification system to
evaluate soil behavioral type (SBT) that uses normalized piezocone readings (Robertson 1990
1991 Lunne et al 1997)
A-10
(119902119905 minus 120590119907119900) (a) normalized tip resistance 119876 = A11 prime 120590119907119900
100∙119891119904 (b) normalized sleeve friction 119865() = A12
(119902119905 minus 120590119907119900)
(1199062 minus 1199060) (c) normalized porewater pressure 119861119902 = A13
(119902119905 minus 120590119907119900)
While the Q-F-Bq classification is a three-dimensional plot it is often presented as a set of two-dimensional
plots specifically Q vs F and Q vs Bq as shown in Figure A6 Data are grouped according to layers and
can be superimposed to identify their association with the nine soil behavioral types All indirect CPT soil
classification approaches should be cross-checked and verified for a particular geologic setting and local
geotechnical conditions before routine use in practice
Figure A6 Nine-zone soil behavioral type charts (a) normalized cone resistance vs normalized
sleeve friction (b) normalized cone resistance vs normalized porewater pressure parameters
(adapted after Robertson 1991 Lunne et al 1997)
The development of a CPT material index (Ic) has been found advantageous in the initial screening of soil
types and helps to organize the sounding into their respective SBT zones of similar soil response In this
case the CPT index is found from (Robertson 2009)
A-11
A2 22 )log221()log473( rtnc FQI
where Qtn = modfied stress-normalized net cone resistance given by
(119902119905 minus 120590119907119900)120590119886119905119898 119876 = A31 prime (120590119907119900120590119886119905119898)119899
where σatm = reference stress equal to atmospheric pressure (1 atm asymp 1 bar = 100 kPa asymp 1 tsf asymp 147 psi)
The exponent n is soil-type dependent n = 1 (clays) n asymp 075 (silts) and n asymp 05 (sands) If units of bars are
used then Qtn (units of bars) is simply
(119902119905 minus 120590119907119900) 119876 = 119899 A32
prime ) (120590119907119900
The operational value of exponent n is found by iteration Initially an exponent n = 1 is used to calculate
the starting value of Ic (ie Qtn = Q) and then the exponent is upgraded to
prime 120590119907119900 119899 = 0381 ∙ 119868119888 + 005 ( ) minus 015 le 10 A4 120590119886119905119898
Then the index Ic is recalculated Iteration converges quickly and generally only 3 cycles are needed to
secure the operational Ic at each depth The soil zones and associated Ic values are detailed in Figure A7
The sensitive soils of zone 1 can be screened using the following expression
A-12
Zone 1 119876119905119899 lt 12119890119909119901( minus14 ∙ 119865119903)
A5
If any soil layers are found within zone 1 (sensitive soils) then caution should be exercised as these
structured clays are prone to instability collapses and difficulties in construction and performance
Another indicator of zone 1 sensitive soils is when the normalized porewater parameter Bq gt 08 When
these criteria are evident the geoengineer should consult with senior geotechnical personnel or the chief
engineer for guidance
Figure A7 Delineation of soil behavioral type zones using CPT material index Ic
A-13
Stiff overconsolidated clayey sands and sandy clays soils of zone 8 (15 lt Fr lt 45) and zone 9 (Fr gt
45) can be identified from the following criterion
Zones 8 and 9 0020)1(00030)1(0050
12
rr
tnFF
Q A6
The remaining soil types are identified by the CPT material index Zone 2 (organic clayey soils Ic ge 360)
Zone 3 (clays to silty clays 295 le Ic lt 360) Zone 4 (silt mixtures 260 le Ic lt 295) Zone 5 (sand mixtures
205 le Ic lt 260) Zone 6 (clean sands 131 le Ic lt 205) and Zone 7 (gravelly to dense sands Ic le 131)
The red dashed line at Ic = 260 represents an approximate boundary separating drained (Ic lt 260) from
undrained behavior (Ic gt 260)
Once the specific zone has been assigned to a layer a visual representation can be made to show
either the zone number or colorization so that the predominant layers and soil types can be
realized
A123 Probabilistic Soil Types from CPT
The inference of soil type has been addressed using probabilistic methods as discussed by Abu-
Farshakh et al (2008) and Tumay et al (2013) Here the CPT readings can be post-processed to
provide the percentage of sand silt and clay components with depth The output is consistent
and compatible with the current MnDOT soil classification chart (Figure A8) which is similar in
content to that used by the USDA
A-14
Figure A8 Chart for soil classification system by MnDOT
A13 Identifying Soft Organic Soils
The identification of soil type using cone penetration tests (CPT) is usually done on the basis of empirical
soil behavioral type (SBT) charts that use the measured readings (qt fs andor u2) While there at least
25 different sets of charts available (ie Hegazy 1998 Eslami et al 2017 Valsson 2016) one of the most
widely used and popular is the 9-zone SBT system that uses three normalized cone readings Qtn Fr and
Bq (Robertson amp Cabal 2007 Lunne et al 1997) The system is denoted as SBTn and plotted on charts in
terms of Q vs F and Q vs Bq as shown in Figure A9
A-15
1
10
100
1000
01 1 10
Norm
aliz
ed
Con
e T
ip R
esis
tan
ce
Q
Normalized Friction Fr = 100 fst(qt - svo) ()
9-Zone Soil Behavioral Type Chart
Gravelly Sands(zone 7)
Sands
(zone 6)
Sandy Mixtures(zone 5)
Silt Mix(zone 4)
Clays
(zone 3)
Organic Soils(zone 2)
Sensitive Clays
and Silts(zone 1)
Very stiff OC clayto silt (zone 9)
Very stiffOC sandto clayey
sand(zone 8)
1
10
100
1000
-06 -04 -02 00 02 04 06 08 10 12 14
No
rmal
ized
Co
ne
Tip
Res
ista
nce
Q
Normalized Porewater Bq = Du2(qt-svo)
Clays(zone 3)
Sensitive(zone 1)Organic
(zone 2)
Gravelly Sands(zone 7)
Sands(zone 6)
Figure A9 CPT charts for SBTn system (a) Q versus F (b) Q versus Bq
Of particular concern in geotechnical site characterization is the proper identification of soft
organic soils such as organic clays and silts (OL and OH) and peats (Pt) These soils often cause
difficulties and problems during construction and performance of civil engineering works because
of bio-degradation long-term creep excessive settlements andor other issues
The SBTn system has a category labelled Zone 2 recognizes organic soils However several recent
studies have noted that data from CPTu soundings do not always fall within the bounds
delineated using Zone 2 even though laboratory tests and field inspections clearly note the
presence of organic soils Lab tests to classify organic soils includes loss on ignition (LOI gt 5)
and two pairs of Atterberg limits testings per ASTM standards as well as the visual-manual
identification of strong organic odor and smell and dark color notably black and dark gray
Results by Mlynarek et al (2014) show that organic soils in Poland are not adequately identified by the
Q-F chart as seen in Figure A10 Moreover studies by Zawrzykraj et al (2017) found that neither the Q-
A-16
F and Q-Bq plots worked well to recognize organic soils and peats in Poland Nejaim et al (2016) found
that Brazilian soft organic clays were misclassified as Zone 3 (clays) rather than Zone 2 (organic clays)
Similarly a recent study on CPTu data from organic clays located at 24 sites in the USA Sweden Mexico
Brazil and Australia have been compiled (Agaiby 2018) These data are shown to generally avoid falling
with the Zone 2 boundaries in either Q-F or Q-Bq charts thus are not recognized during these soundings
as indicated by Figure A11 The exception in this case is the Mexico City clay which is correctly identified
however the other 23 sites are generally not properly recognized and categorized Additional findings by
Missiaen et al (2015) found that the CPTu results in Belgian soils also did not identify organic clays and
peats in the non-normalized version of these charts that uses Qt versus friction ratio (Rf = 100middotfsqt) as
detailed by Robertson amp Cabal (2007)
Figure A10 Soil behavioral chart with superimposed CPTu data from Polish organic soils
(from Mlynarek et al 2014)
A-17
Figure A11 Paired SBTn charts with CPTu data from 24 organic clay sites indicating non-compliance
with Zone 2 (organic soils)
(compiled by Agaiby 2018)
A14 Simplified Yield Stress Evaluation in Clays by CPTu
From the development of a hybrid formulation of spherical cavity expansion (SCE) theory with critical-
state soil mechanics (CSSM) the effective preconsolidation stress or yield stress (σp) of clays can be
expressed in terms of three piezocone parameters (a) net cone tip resistance qnet = qt - σvo (b)
measured excess porewater pressure Δu2 = u2 - u0 and (c) effective cone resistance qE = qt - u2 Details
on the SCE-CSSM solutions are given elsewhere (Mayne 1991 Mayne 2017 Agaiby amp Mayne 2018)
A-18
For normal and well-behaved clays which include inorganic fine-grained soils such as clays and silts
of low sensitivity a set of simple relationships can be derived By adopting characteristic default values
for effective friction angle ϕ = 30deg and rigidity index (IR = 100) linear expressions for the effective
preconsolidation stress are obtained
prime 120590119901 = A7 033 ∙ 119902119899119890119905
prime 120590119901 = 053 ∙ ∆1199062 A8
prime 120590119901 = 060 ∙ 119902119864 A9
A141 Application to soft Chicago clay
An example of a common geomaterial in this category includes the infamous soft Chicago clays
that were deposited in a glacio-lacustrine environment (Cho amp Finno 2010) The soft clay has a
mean water content of 20 liquid limit of 38 and plasticity index PI= 12 Results of CPTu tests
conducted at the national geotechnical experimentation site (NGES) at Northwestern University
are shown in Figure A12 (Ouyang amp Mayne 2017) As seen in the profile a thin soft clay resides
between depths of 9 to 105 m with a thicker soft clay layer in the range of 12 to 18 m
A-19
0
2
4
6
8
10
12
14
16
18
20
00 02 04 06 08 10 12D
ep
th (
m)
CPTu qt and u2 (kPa)
sand (SP)
Blodgett
Park Ridge
Deerfieldsoft clay
ClayLayerof Study
Northwestern UnivNational Test Site
Figure A12 Results of CPTu sounding at the NGES at Northwestern University Illinois
Post-processing of the CPTu data using Equations A7 A8 and A9 show good agreement in evaluating the
profile of effective yield stress in this clay layer compared with results from one-dimensional
consolidation tests made on undisturbed samples taken at the site
A-20
10
11
12
13
14
15
16
17
18
19
20
0 100 200 300 400 500 600 700 800 900 1000
De
pth
(m
)Effective Yield Stress sp (kPa)
Consols
033 qnet
053 Du2
060 qE
svo
Figure A13 Evaluation of effective yield stresses at NWU using consolidation tests and CPTu
A142 Application to soft Bay Mud California
Another example of a well-behaved fine-grained soil is the San Francisco Bay Mud which is a soft clay
deposit in northern California Results of a representative CPTu are shown in Figure A14 and indicate the
soil profile at a site in eastern San Francisco (Hunt et al 2002) For this site index values include 70 lt
wn lt 90 60 lt LL lt 80 and 35 lt PI lt 45 Post-processing the CPTu data using Equations A7 A8 and
A9 show good agreement with each other as well as with the results of consolidation tests as seen in
A-21
Figure A15 The consolidation tests were performed using a CRS device as reported by Pestana et al
(2002)
San Francisco Bay Mud (Pestana et al 2002)
0
5
10
15
20
25
30
35
40
0 1000 2000 3000
De
pth
(m
ete
rs)
Cone Resistance qt (kPa)
0
5
10
15
20
25
30
35
40
0 10 20 30 40 50 60
Sleeve Friction fs (kPa)
0
5
10
15
20
25
30
35
40
0 500 1000 1500
Porewater u2 (kPa)
Miscellaneous Fill
Young Bay Mud
Young Bay Mud
Clayey Sand
Miscellaneous Fill
Young Bay Mud
Young Bay Mud
Figure A14 Representative CPTu in San Francisco Bay Mud (data from Hunt et al 2002)
A-22
San Francisco Bay Mud (Pestana et al 2002)
0
5
10
15
20
25
30
0 100 200 300 400 500 600
De
pth
(m
ete
rs)
Preconsolidation Stress sp (kPa)
svo
033 qnet
053 Du2
060 qE
CRS
svo
Figure A15 Evaluation of effective yield stresses in Bay Mud from consolidation tests and CPTu
A15 CPTu Screening of Soft Organic Clays
For soft organic clays the Equations A7 A8 and A8 will not apply thus can serve as a means to
screen such geomaterials from the soil profile Two examples of CPTu soundings in soft organic
clays are presented to illustrate the approach Additional case records and documentation of
CPTu data in organic clays are given in Agaiby (2018)
A151 Soft organic alluvial soils in Washington DC
Results of in-situ tests including CPTu soundings in soft organic clayey silts along the Potomac River at the
Anacostia Naval Air Station are presented by Mayne (1987) Here the soft soils are categorized as organic
clayey silts (OH) per the Unified Soils Classification Systems (USCS) with mean values (and plus and minus
A-23
one standard deviations) of wn = 68 plusmn 16 LL = 83 plusmn 25 and PI = 37 plusmn 17 A representative
piezocone sounding is depicted in Figure A16 For the post-processing of CPTu data the use of the Q-F
and Q-Bq graphs do not find any points within the Zone 2 category for organic soils When the data are
evaluated to determine the profile of effective yield stress the Equations A7 A8 and A9 do not agree as
shown by Figure A17
0
5
10
15
20
25
0 2000 4000 6000
Dep
th (
met
ers)
Cone Resistance qt (kPa)
0
5
10
15
20
25
0 20 40 60 80 100
Sleeve Friction fs (kPa)
0
5
10
15
20
25
0 200 400 600
Pore Pressure u2 (kPa)
Figure A16 Representative CPTu in soft organic clay along the Potomac River DC
A-24
0
5
10
15
20
25
0 3 6 9D
epth
(m
eter
s)Soil Behaviour Type Zone
Q and Fchart
0
5
10
15
20
25
0 100 200 300 400 500 600
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
033qnet
053 Du2
06 qE
Oed
Figure A17 Post-processing of CPTu data from Anacostia-Bolling (a) SBTn zones using CPT material
index Ic (b) CPTu screening equations for yield stress
Note that the hierarchy of the results from Equations A7 A8 and A9 show that the CPTu
evaluations for preconsolidation stress in soft organic clays indicates
[A10] σp = 053 Δu2 lt 033 qnet lt 060 qE
A152 Soft organic peaty clay in Saint Paul MN
Results of CPTu tests were collected during a training exercise underneath I-35E just northeast
of Saint Paul MN in 2007 All three MnDOT CPT rigs available at the time were used to obtain
in-situ test data at the site The site is well known as having soft organic soils and samples were
obtained from three borings made at the site (Boring ID T523 T542 and T518) Boring logs
A-25
indicate the presence of highly organic silt loams with marl and spongy organic clayey silts
From 12 samples the natural water contents ranged from 30 to 191 with a mean of 112 plusmn
58
A representative piezocone sounding from the site (MnDOT CPTu ID F22Y0703C) is used for
this illustration and is presented as Figure A18 Post-processing these data using the SBTn
algorithms and CPT material index show that the soils classify primarily as Zone 3 (clays) and
Zone 4 (silty mix) thus do not identify the geomaterials properly as Zone 2 (organic soils)
0
5
10
15
0 500 1000 1500 2000
Dep
th (
met
ers)
Cone Tip Resistance
qt (kPa)
0
5
10
15
0 10 20 30 40 50 60
Sleeve Friction
fs (kPa)
0
5
10
15
-100 0 100 200 300 400
Porewater Pressure
u2 (kPa)
Figure A18 MnDOT CPTu sounding at the I-35E test site near St Paul Minnesota
A-26
1
10
100
1000
01 1 10
Norm
aliz
ed
Con
e T
ip R
esis
tan
ce
Q
Normalized Friction Fr = 100 fst(qt - svo) ()
9-Zone Soil Behavioral Type Chart
St PaulGravelly Sands(zone 7)
Sands
(zone 6)
Sandy Mixtures(zone 5)
Silt Mix(zone 4)
Clays
(zone 3)
Organic Soils(zone 2)
Sensitive Clays
and Silts(zone 1)
Very stiff OC clayto silt (zone 9)
Very stiffOC sandto clayey
sand(zone 8)
1
10
100
1000
-06 -04 -02 00 02 04 06 08 10 12 14
No
rmal
ized
Co
ne
Tip
Res
ista
nce
Q
Normalized Porewater Bq = Du2(qt-svo)
Clays(zone 3)
Sensitive(zone 1)Organic
(zone 2)
Gravelly Sands(zone 7)
Sands(zone 6)
Figure A19 Post-processing St Paul sounding using the 9-zone SBTn classification system
Using the recommended approach Figure A20 shows the CPT-evaluated yield stress profiles from
equations A7 A8 and A9 Here again the three estimates of σp do not agree but show the same
hierarchy as detailed in Equation A10
A-27
0
5
10
15
20
0 50 100 150 200 250 300 350 400
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
06 qeff svo
033qnet Consols
054 Du2
Figure A20 CPTu screening of soft organic soils at St Paul test site
A16 CPTu Evaluations of Yield Stress in Soft Organic Soils
Once the presence of soft organic clays and silts is identified using the aforementioned CPT
screening algorithms the evaluation of effective yield stress is conducted using the procedure
outlined in Chapter 2 of this MnDOT manual For soft organic soils an exponent of m = 090 is
recommended (Mayne et al 2009) Therefore the preconsolidation stress is obtained from
(units of kPa)
prime = A11 120590119901 033 ∙ (119902119899119890119905)090
The algorithm can be expressed in dimensionless form by
prime = A12 120590119901 033 ∙ (119902119905 minus 120590119907119900)119898prime ∙ (120590119886119905119898100)1minus119898prime
A-28
where σatm = reference stress (= 100 kPa = 147 psi) and m = 090 for soft organic silts and clays
The approach is applied to the two former example case studies in Figure A21 (Washington DC)
and Figure A22 (St Paul MN) respectively showing good agreement with benchmark results
taken from laboratory consolidation tests on undisturbed samples of these sites in both cases
0
5
10
15
20
25
0 100 200 300 400 500 600
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
033qnet^090
Oed
Figure A21 Profiles of yield stress from consolidation tests and CPTu method for organic clayey silts at
Anacostia-Bolling site in Washington DC
A-29
0
5
10
15
20
0 20 40 60 80 100 120 140 160
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
CPT
Consols
120590119901prime = 033 119902119899119890119905
090 in kPa
Figure A22 Profiles of yield stress from consolidation tests and CPTu method for organic clayey silts at
MnDOT test site near Saint Paul Minnesota
A17 Soil Unit Weight
As shown by Figure A23 total soil unit weight (t) can be estimated from the CPT sleeve friction resistance
(Mayne 2014) The equation can be expressed by
120574119905 = 120574119908 ∙ [122 + 015 ∙ 119897119899 (100 lowast 119891119904120590119886119905119898 + 001)] A13
where w = unit weight of water and satm = atmospheric pressure
A-30
Figure A23 Evaluation of soil unit weight from CPT sleeve friction
A18 Effective Stress Friction Angle
Sands
The effective stress friction angle () is one of the most important soil properties as it governs the strength
of geomaterials as well as affects soil-pile interface and pile side friction While an effective cohesion
intercept (c) can also be considered this is usually reserved for cemented or bonded geomaterials or
unsaturated soils and may lose its magnitude with time ageing or with prolonged environmental
changes
For clean quartz to silica sands where porewater pressures are essentially hydrostatic (Bq = 0) the
following expression has been calibrated with triaxial compression test results from undisturbed sand
samples and normalized cone resistances as presented in Figure A24 (Mayne 2007 2014)
A-31
120601prime(119889119890119892) = 176deg + 110deg 119897119900119892 (1199021199051) A14
Figure A24 Evaluation of effective stress friction angle in quartz-silica sands from CPT
Note Relationship applies to drained soil behavior when Ic lt 26 andor Bq lt 01
where qt1 is an earlier form of stress normalized cone tip resistance for sands given by
(119902119905 120590119886119905119898) 1199021199051 = A151 05 prime (120590119907119900120590119886119905119898)
A-32
In terms of units of bars this is expressed simply as (in units of bars)
119902119905 1199021199051 = 05 A152 prime ) (120590119907119900
Recently Robertson amp Cabal (2015) recommended the use of the modified form of normalized cone
resistance (Qtn) in Equation A10
120601prime (119889119890119892) = 176deg + 110deg 119897119900119892 (119876119905119899) A16
In sands qnet asymp qt since the overburden term is small relative to the cone tip resistance Figure A25 shows
that the two normalizations give very comparable results
Figure A25 Comparable values from qt1 and Qtn normalization schemes for CPTs in sands
Clays and Silts
A-33
In the case of soft to firm intact clays and silty clays the effective friction angles are determined from the
normalized cone resistance and porewater pressure parameters (Senneset et al 1989 Mayne 2016) as
shown in Figure A26 The exact solution when the angle of plastification = 0 is given as
qBQ
)tan1(tan61
1)tanexp()245(tan 2
A17
Approximate NTH Solution for from CPTu
qBQ
)tan1(tan61
1)tanexp()245(tan2
Approx asymp 295deg∙Bq0121∙[0256 + 0336∙Bq + log Q ]
1
10
100
20 25 30 35 40 45
Co
ne
Res
ista
nce
Nu
mb
er Q
(degrees)
Bq
010204060810
c = 0 = 0deg
Theory (dots)
Approximation (lines)
Theory
Figure A26 Evaluation of effective stress friction angle in clays and silts from CPTu results
Note Relationship applies to undrained soil behavior when Ic ge 26 andor Bq ge 01
which can be approximately inverted into the form (Mayne 2007)
A-34
0121 120601prime = 295deg ∙ 119861119902 ∙ [ 0256 + 0336 ∙ 119861119902 + log(119876)] A18
This algorithm is specifically applicable for the following ranges of porewater pressure parameter (01 le
Bq lt 1) and effective stress friction angles (20deg le lt 45deg)
A19 Stress History
The stress history can be characterized by an apparent yield stress or preconsolidation stress (sp) as well
as by its normalized and dimensionless form YSR = spsvo = yield stress ratio The YSR is in effect the
same as overconsolidation ratio (OCR) however now generalized to accommodate mechanisms of
preconsolidation that occur beyond just erosion glaciation and removal of overburden stresses but also
due to ageing desiccation repeated cycles of wetting-drying bonding repeated freeze-thaw cycles
groundwater changes and other factors
A-35
A generalized approach for evaluating the yield stress or preconsolidation in soils using net cone
resistance has been formulated as presented in Figure A27 (Mayne 2015) Yield stress in soils from CPT
10
100
1000
10000
10 100 1000 10000 100000
Yie
ld S
tre
ss
s
y
(kP
a)
Net Cone Resistance qt - svo (kPa)
Blessington
General Trend
sy = 033(qt-svo)m
Fissured Clays m = 11
Intact clays m = 10 Sensitive clays m = 09
Silt Mixtures m = 085Silty Sands m = 080
Clean Sands m = 072
m = 11 10 09
085
080
072
Units kPa
Fissured Clays
Figure A27 Evaluation of yield stress or preconsolidation stress in soils from CPT
The algorithm can be expressed in dimensionless form by
1minus119898prime prime 120590119886119905119898 120590119901 = 033 ∙ (119902119905 minus 120590119907119900)119898prime
( ) A19 100
where m = exponent depends on soil type m = 1 (intact clays) 085 (silts) 080 (sandy silts to silty sands)
and 072 (sands) For fissured clays the exponent m may be 11 or higher depending upon the age of the
A-36
formation and degree of jointing and discontinuities Note that fissured clays can be identified when Ic lt
26 and Bq lt 01
The exponent m has also been calibrated with CPT material index as presented in Figure A28
An algorithm to express this relationship for non-fissured soils is given by
25)652(1
2801
cIm
A20
This allows the post-processing of CPT to automatically choose the appropriate exponent m for
a layer by layer analysis
Figure A28 Yield stress exponent m in terms of CPT material index Ic
A-37
A110 Lateral Stress Coefficient
The horizontal geostatic state of stress is represented by the lateral stress coefficient K0 = shosvo
commonly referred to as the at-rest condition The magnitude of K0 for soils that have been loaded and
unloaded can be approximately estimated from
119870119900 = (1 minus 119904119894119899120601prime) ∙ 119874119862119877119904119894119899120601prime A21
Data compiled from in-situ K0 measurements using self-boring pressuremeter tests (SBPMT) total stress
cells (TSC) or push-in spade cells andor laboratory methods (instrumented consolidometers triaxials
andor suction measurements) on a variety of clays silts and sands have been compiled and reported by
Ku amp Mayne (2015) as shown in Figure A29 verifying Equation A15 as a means for evaluating K0 in soils
Figure A29 Relationship between lateral stress coefficient K0 and YSR or OCR in soils (Ku amp Mayne
2015)
A-38
A111 Undrained Shear Strength
The loading of soils can occur under conditions of being fully drained (Du = 0) partially drained or full
undrained (DVV0) where Du = excess porewater pressures (above hydrostatic) and DVV0 = volumetric
strain Further details on specific stress paths are best explained in terms of critical state soil mechanics
or CSSM (Mayne et al 2009 Holtz Kovacs amp Sheahan 2011) The prevailing drainage conditions depend
upon the rate of loading and permeability characteristics of the soil Normally in sands that are pervious
and exhibit high permeability a drained response occurs Exception to this may occur in loose sands
during fast earthquake loading resulting in soil liquefaction In clays that exhibit low permeability a fast
rate of loading will result in undrained loading at constant volume This in fact is a temporary and
transient condition often termed short term loading In the long term eventually porewater pressures
will dissipate (albeit slowly) and a drained response will prevail thus termed long term loading
The overall soil strength is controlled by the effective stress strength envelope Most commonly this is
represented by a simple linear relationship termed the Mohr-Coulomb criterion where the maximum
shear stress (max called the shear strength) is given as
= 119888prime + 120590prime ∙ 119905119886119899 120601prime A22 120591119898119886119909
where c = effective cohesion intercept s = effective normal stress and = effective stress friction angle
As a starting point values of c = 0 and = 30deg can be adopted for all soil types at least until the specific
soil type geologic formation and results from high-quality laboratory or field test data are available
Peak Undrained Shear Strength from Stress History
Using simplified critical state soil mechanics the peak undrained shear strength (max = su) can be
evaluated from the effective stress strength envelope (c = 0 effective friction angle ) and stress history
(ie OCR) in the form
prime DSS 119904119906 = (119904119894119899120601prime2) ∙ (119874119862119877)Λ ∙ 120590119907119900 A23
A-39
which corresponds to a simple shear mode Figure A30 presents the aforementioned relationship
together with data from 17 different clays tested in simple shear
Figure A30 Normalized undrained shear strength from simple shear tests on various clays
versus YSR or OCR
For the triaxial compression (TC) mode the equation would give slightly higher strengths that are
calculated from
prime TC 119904119906 = (1198721198882) ∙ (1198741198621198772)Λ ∙ 120590119907119900 A24
where Mc = (6 middot sin)(3-sin)
Peak Undrained Shear Strength from CPTu
A more direct approach to assessing undrained shear strength is via bearing capacity theory
whereby
A-40
A25 kt
votu
N
qs
s
where Nkt = bearing factor that depends on mode of shearing sensitivity of the clay degree of
overconsolidation and other factors For soft offshore clays the back figured Nkt from data collected at
14 well-documented sites (Low et al 2010) determined mean values based on mode of shearing Nkt =
119 (triaxial compression) Nkt = 136 (simple shear) and Nkt = 133 (vane)
A recent study of 51 clays that were tested by both field piezocone and laboratory CAUC triaxial tests
showed that essentially Nkt = 12 for intact clays and clayey silts of low to medium sensitivity (Mayne
Peuchen amp Baltoukas 2015) Figure A31 shows the slightly different trends for offshore versus onshore
clays For sensitive clays a lower value of Nkt = 10 would be appropriate and for fissured clays a higher
value of Nkt would be in the range of 20 to 30
A-41
A-42
Figure A31 Database from 51 clays with CPT qnet versus lab measured triaxial shear strength
(Mayne Peuchen amp Baltoukas 2015)
For soft to firm clays an alternate means to evaluate undrained shear strength is via the excess porewater
pressures ( u = u2 - u0) from the following
u
uN
us
D
D 2
A26
where N u = porewater bearing factor also dependent on the aforementioned factors The study by
Low et al (2010) found N u = 59 (triaxial compression) N u = 69 (simple shear) and N u = 71 (vane
shear)
A third alternate is found from the effective cone resistance (qE = qt - u2) whereby
kE
tu
N
uqs 2
A27
where NkE is a bearing term for this approach Mayne et al (2015) indicated a mean value of NkE = 8 for
soft-firm intact clays
Remolded Undrained Shear Strength from CPT
The remolded undrained shear strength (sur) is obtained from either field vane tests lab mini-vane shear
tests or lab fall cone devices This affords the evaluation of the clay sensitivity (St) which is defined as the
ratio of peak to remolded strengths at a given water content
119904119906(119901119890119886119896) 119878119905 = frasl A28 119904119906119903
For the CPT the sleeve friction has been noted to give values that are comparable to the
remolded undrained shear strength (Powell amp Lunne 2015) Therefore
asymp 119891119904 A29 119904119906119903
A112 Ground Stiffness and Soil Moduli
The stiffness of the ground can be represented by several geoparameters including the compressibility
indices (Cr Cc Cs) spring constants (kv) and moduli Regarding the latter there are several moduli that
are used in geotechnical engineering including shear modulus (G and Gu) Youngs modulus (E and Eu)
constrained modulus (D) bulk modulus (K) resilient modulus (MR) and subgrade reaction modulus (ks)
A-43
Soil Modulus
The definition of modulus is taken as E = DsD with Figure A32 showing a typical deviator stress (q = s1
- s3) versus axial strain (a) curve Theoretical interrelationships between the elastic moduli G E D and
K have a dependence on the Poissons ratio v The value of Poissons ratio can be taken as vu = 05 for
undrained loading (ie constant volume) while for drained loading which is accompanied by volumetric
strains a value of v = 02 may be used If we adopt the reference modulus as E then the
interrelationships with the other elastic moduli are given by
a Reference Stiffness 119864prime = drained Youngs modulus A30
119864prime
b Shear Modulus 119866prime = A31 [2(1+120584prime)]
119864prime (1minus120584prime) c Constrained Modulus 119863prime = A32
[(1+120584prime)(1minus2120584prime)]
119864prime
d Bulk Modulus 119870prime = A33 [3middot(1minus2120584prime)]
A-44
Figure A32 Representative stress-strain-strength curve for soils in triaxial compression
For the extreme case when = 0 in fact D = E When = 02 the two moduli are only 10
different D = 11middotE thus the constrained modulus and drained Youngs modulus are often
considered somewhat interchangeable
The resilient modulus (MR) applies to pavement analysis and design most commonly measured by cyclic
triaxial testing under repeated load applications In fact MR is a special case of Youngs modulus that
relates to the small strain stiffness measured in the nondestructive range but has developed permanent
plastic strains after many cycles of loading (Brown 1996 Dehler amp Labuz 2007)
The subgrade reaction modulus is actually a combined soil-structural parameter as its value
depends on the ground stiffness and the size of the loaded element The subgrade modulus is
defined as ks = q where q = applied stress and = measured deflection In terms of elasticity
solutions the deflection of a flexible circle of diameter d is given by = qmiddotdmiddot(1-2)E Therefore
119864prime
119896119904 = A34 [119889middot(1minus1205842)]
which has units of kNm3 or pcf
Table A2 summarizes several selected studies towards the approximate evaluation of these moduli
obtained directly from measured CPT data Note however that the results from in-situ penetrometer
tests generally represent a peak strength as indicated in Figure A32 Thus the measured cone tip
resistance (qt) reflects the top of the stress-strain curve either the undrained strength in clays (su) or the
effective friction angle () of sands or even a strength intermediate between these two values Thus
A-45
Source Soils Studied Reference Modulus
Findings
Kulhawy amp Mayne 1990
8 stiff OC clays Lab constrained modulus
D asymp 825 middot (qt - svo)
Mohammed et al (2000)
Resilient modulus
Abu-Farsakh 2004
7 Louisiana sites
Lab constrained modulus
D asymp 358 middot (qt - svo)
Mayne (2007b)
All soil types Constrained moduli from lab consolidation
First-order estimate D asymp middot(qt - svo)
Robertson (2009)
All soil types Constrained modulus from consolidation
D = m middot (qt - svo)
when Ic gt 22
1 m = Qtn when Qtn le 14
2 m = 14 when Qtn gt 14
when Ic lt 22 then
= 003 middot 10055 Ic + 168 m
Liu et al (2016)
16 clay sites in China
Resilient modulus MR = (146qt 053 + 1355 fs
14 + 236)244
(all in MPa)
Casey et al (2016)
8 clays Triaxial tests for undrained Youngs modulus
Eu(svc)07 = 316 - 23 middot LL
where Eu and svc (MPa)
and LL = liquid limit ()
qt is probably not the best means to obtain the slope of the stress-strain curve Instead the SCPTu that
provides the initial tangent shear modulus (Gmax) from the shear wave velocity (Vs) is a better choice
Table A2 Selected Modulus Relationships with Cone Penetration Test Measurements
A-46
Nonlinear Modulus
The small-strain shear modulus (Gmax or G0) represents the initial stiffness of all soils and rocks It is the
beginning of all stress-strain-strength curves for geomaterials and obtained from elasticity theory
119866119898119886119909 = 1198660 = 120588119905 middot 1198811199042 A35
where Vs = shear wave velocity t = tga = total soil mass density t = total soil unit weight and ga =
gravitational acceleration constant (98 ms2)
From a general viewpoint on stiffness the shear modulus G of soil is defined as the slope of shear stress
versus shear strain G = DDs for a tangent definition and by G = s as a secant definition The shear
modulus is related to its associated Youngs modulus E = 2G(1+v) Both moduli are in fact highly
nonlinear ranging from a maximum value at the small-strain stiffness (Gmax = G0 = t middotVs2 where t =
total soil mass density and Vs = shear wave velocity) to intermediate G values at medium strains (asymp 1)
to low values at peak strength As such a variety of algorithms and formulae have been developed to
represent either a partial range or the full stress-strain-strength behavior of soils over a range of
interests (eg Mayne 2005) In these formulations a variable number of input parameters may be
required in order to produce a stress-strain-strength curve
A modified hyperbolic form suggested by Fahey amp Carter (1993) has favorable attributes in that the
modulus reduction factor can be established with only a single variable This allows the initial stiffness
to the be small-strain stiffness (G0) that is reduced in terms of level of mobilized strength eg max =
qqmax = 1FS which is simply the reciprocal of the calculated factor of safety (FS) The magnitude of
secant shear modulus (G) corresponding to the particular level of loading is given by
119866 = 119872119877119865 middot 1198660 = (1198661198660) middot 1198660 A36
A-47
where the MRF = modulus reduction factor determined as
1 minus (120591 )119892 119872119877119865 = (1198661198660) = frasl A37 120591119898119886119909
and g = empirical exponent term Figure A33 shows the relationships for normalized shear stress ()
versus shear strain (s) and corresponding normalized secant shear modulus (G = s) with shear strain
over a range of exponent values 01 le g le 10 For a discussion of tangent G please see Fahey amp Carter
(1993)
Using laboratory data from resonant column-torsional shear tests and triaxial specimens with local
strain measurements for Gmax reference values modulus reduction curves for a selection of sands and
clays are presented in Figure A34 The data indicate that the exponent g falls within the range 02 lt g lt
05 for many soils tested drained and undrained A mean value of g = 03 is recommended for
preliminary site investigations and designs until additional information can be obtained
The value of max is the shear strength of the soil generally taken as either (1) drained (c = 0) max =
svo middot tan or (2) undrained where max = su = undrained shear strength as discussed earlier Thus each
shear stress () can be associated with its shear modulus (G) and the relevant shear strain is found from
s = G This allows for generation of nonlinear stress-strain-strength curves at all depths from SCPTu
data in clays sands and mixed soil types
A-48
Modified Hyperbola (Fahey amp Carter 1993 Canadian Geotech J) Coefficient Term f = 1
``
0
01
02
03
04
05
06
07
08
09
1
000001 00001 0001 001 01 1
No
rmal
ized
GG
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
0
01
02
03
04
05
06
07
08
09
1
0 01 02 03 04 05 06 07 08 09 1
No
rmali
zed
GG
ma
x
Mobilized Shear Stress max
g =1
07
05
03
02
01
0
01
02
03
04
05
06
07
08
09
1
00001 0001 001 01 1
No
rmali
zed
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
0
01
02
03
04
05
06
07
08
09
1
0 1 2 3 4 5 6 7 8 9 10
No
rmali
zed
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
Reduction Factor
GGmax = 1- (max)g
Figure A33 Normalized modulus (GGmax) and normalized shear stress (max) versus shear strain (s)
for modified hyperbolic relationship of Fahey amp Carter (1993)
A-49
0
01
02
03
04
05
06
07
08
09
1
0 01 02 03 04 05 06 07 08 09 1
Mo
du
lus
Re
du
cti
on
G
Gm
ax
or
EE
ma
x
Mobilized Strength max or qqmax
NC SLB Sand
OC SLB Sand
Hamaoka Sand (e=0625)
Hamaoka Sand
Toyoura Sand e = 067
Toyoura Sand e = 083
Ham River Sand
Ticino Sand
Kentucky Clayey Sand
Kaolin
Fujinomori Clay
Pietrafitta Clay
London Clay
Vallericca Clay
Pisa Clay
Pisa Clay
Pisa Clay
Pisa Clay
Kiyohoro Silty Clay
Thanet Clay
Exp g = 02
Exp g = 03
Exp g = 04
Exp g = 05
MRF = 1 - (qqmax)g
Figure A34 Modulus Reduction Factors (MRF) with mobilized strength of clays and sands
A113 Coefficient of Consolidation
The rate at which foundation and embankment settlements occur as well as the dissipation of excess
porewater pressures is controlled by the coefficient of consolidation (cv) The magnitude of cv is also
required in the design of vertical wick drains that can be installed in soft ground to expedite the time for
consolidation Using the results of CPTu dissipation tests which measure the rate at which the u2 readings
vary with time the in-situ profile of cv can be evaluated Often the piezocone-interpreted values of cv
are validated by comparison with results from laboratory one-dimensional consolidation tests (eg Abu-
Farsakh MY 2004) Alternatively a better method is to cross-check the values with the measured full-
scale performance of instrumented embankments that are constructed over the ground and the recorded
time-rate of consolidation and settlements can provide the best cv for that site (Abu-Farsakh MY et al
2011)
A-50
Monotonic Dissipation Tests
An illustrative dissipation curve from piezocone testing at IDT State Highway 95 at an embankment and
bridge crossing is presented in Figure A35 After the CPTu sounding was halted at a depth of 512 feet
the recorded decay of porewater pressures was observed to be monotonic with time until the dissipations
were ended at 1000 seconds During that time the u2 readings decreased from 115 psi to 47 psi At this
location the depth to the groundwater table is about 16 feet thus the calculated equilibrium water
pressure is 15 psi Commonly a characteristic time for piezo-dissipations is taken at 50 degree of
consolidation although other degrees may be adopted By evaluating the value of u2 at 50 as shown in
Figure A35 the characteristic t50 = 404 s can be obtained
0
20
40
60
80
100
120
1 10 100 1000 10000
Pore
wat
er P
ress
ure
u2
(psi
)
Time (s)
CPTU-02E-1 z = 1560 m
uo (hydrostatic)
u2 (initial) = 115 psi
u2 (final projected) = u0 = 15 psi
u2 at 50 = frac12(115+15) = 65 psi
Idaho DOT State Route 95Dissipation at 512 feet
t50 = 404 seconds
Time to reach50 consolidation
A-51
Figure A35 Piezo-dissipation at Sandpoint Bridge Crossing Idaho for depth z = 512 feet illustrating the
evaluation of characteristic time t50
It is also common to plot normalized excess porewater pressures relative to the measured initial Du2
that is obtained during penetration at the constant rate of 20 mms ie DuDui versus time Figure A36
shows a set of piezo-dissipation records for a bridge and embankment site in southern Louisiana
reported by the LTRC While the porewater pressure axis is arithmetic the time scale is often plotted on
either logarithmic or square root scales In either case the t50 is then found from the measured
dissipation curve when DuDui = 050
A-52
Figure A36 Set of monotonic piezo-dissipations taken at various depths for Courtableau Bridge
Site on Louisana State Highway 103 (Abu-Farsakh et al 2011)
Dilatory Dissipation Curves
In a number of situations a monotonic type dissipation does not occur on the onset but instead a dilatory
type curve is observed Here after stopping the push of the penetrometer the recorded porewater
pressures initially begin to rise and eventually reach a peak value then afterwards follow a phase where
the Du values decrease to hydrostatic (Figure A37) A full solution is available for both cases (Burns amp
Mayne 2002) Dilatory responses are most often associated with overconsolidated clays and silts
although occasionally are observed in fine-grained soils with low OCRs
The selection of the characteristic t50 value may found by using a square root of time plot to find the initial
u2 value by projection of the post-peak dissipatory data back to the ordinate axis as shown by the example
in Figure A38 In this example the initial u2 = 480 kPa and then the same procedures in Figure A35may
apply
A-53
Figure A37 Monotonic and dilatory porewater dissipation responses in soils
(after Burns amp Mayne 1998)
Figure A38 Method for obtaining t50 from dilatory dissipation curves
(Schneider amp Hotstream 2010)
A-54
Interpretation of Dissipation Test Data
There are a number of different solutions available for the interpretation of the coefficient of
consolidation (cv) from the piezocone dissipation curves For monotonic type response the strain path
method (SPM) developed at Oxford University (Teh amp Houlsby 1991) is well-recognized while the Georgia
Tech solution by Burns amp Mayne (2002) is based on spherical cavity expansion and critical state soil
mechanics (SCE-CSSM) and handles both monotonic and dilatory curves For both solutions a simplified
procedure can be recommended that relies on the aforementioned t50 value obtained from the measured
field data as given in Table A3 The units for cv are in cm2s as shown or equivalent
Table A3 Recommended procedures for calculating cv from t50 obtained in dissipation tests
Method Equation for coefficient of
consolidation cv
RemarksNotes Eqn
No
SPM
(Teh amp
Houlsby
1991) 50
2)(2450
t
Iac
Rc
v
t50 = measured time to reach 50
dissipation
IR = Gsu = undrained rigidity index
G = shear modulus
su = undrained shear strength
ac = penetrometer radius (ac = 178
cm for 10-cm2 cone ac = 220 for a
15-cm2 size)
A32
SCE-CSSM
(Burns amp
Mayne 2002) 50
7502 )()(0300
t
Iac Rc
v
A33
A-55
Evaluating rigidity index
Both the SPM and SCE-CSSM solutions require an estimate of the in-situ rigidity index (IR = Gsu) of the
soil If the results of SCPTU are available Krage et al (2014) have derived an expression for IR that
depends upon the small-strain shear modulus net cone tip resistance and effective overburden stress
250750
max50
8111
vonet
Rq
GI
s A38
where consistent units are input for Gmax qnet and svo The value of Gmax is obtained from B29 using the
measured shear wave velocity (Vs) and unit weight evaluated from B7
If the shear wave velocity is not measured it can be estimated from the CPT data A number of
methods have been reviewed for CALTRANS by Wair et al (2012) including one that is applicable
to sands silts and clays of low sensitivity that are inorganic and uncemented
119881 (119898119904) = [101 middot 119897119900119892(119902119905) minus 114]167 middot (100 middot 119891119904119902119905)03 A39 119904
where qt is in units of kPa (Hegazy amp Mayne 1995) An alternative relationship is given by Robertson amp
Cabal (2015)
A114 Hydraulic Conductivity
The hydraulic conductivity is a parameter that expresses the flow characteristics of the soil In
geotechnics it is also called the coefficient of permeability (k) and has units of cms or feetday
Through consolidation theory the hydraulic conductivity relates directly to the coefficient of
consolidation
A-56
A40 D
ck wv
where w = unit weight of water and D = constrained modulus
Therefore one approach to evaluating k can be from the site-specific cv obtained from the
dissipation tests using an estimate of D from one of the relationships given in Table A2
An approach developed for soft normally-consolidated soils that uses t50 directly to assess the magnitude
of k is presented in
Figure A39 (Parez and Fauriel 1988) The mean trendline through the wedges of soil type can be
approximated by
251
50 (sec)251
1)(
tscmkh
A41
A-57
Figure A39 Hydraulic conductivity versus dissipation time for 50 consolidation
(Parez amp Fauriel 1988)
A-58
APPENDIX B
List of Figures
Figure B1 Conventional methods vs direct CPT approach to shallow foundation response B-3
Figure B2 Direct bearing capacity relationship for embedded square and strip footings situated on clean
sands (after Schmertmann 1978) B-6
Figure B3 Direct CPT bearing factor Rk as function of footing width B and embedment depth from finite
element analyses (Tand et al 1995) B-6
Figure B4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio and sand
consistency (Eslaamizaad amp Robertson 1996) B-7
Figure B5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp Salgado (2005) in
terms of footing size relative density and base settlement B-8
Figure B6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and normalized
cone tip resistance (after Eslami amp Gholami 2005) B-8
Figure B7 Direct relationship between ultimate bearing stress in clays and measured cone tip resistance
(Schmertmann 1978) B-10
Figure B8 Direct relationship between ultimate foundation bearing stress in clays and net cone tip
resistance (after Tand et al 1986) B-11
Figure B9 Measured load-displacements for each of the five large footings at TAMU (Briaud amp Gibbens
1994) sand site B-12
Figure B10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU (Briaud amp
Gibbens 1994) footings B-13
Figure B11 Characteristic stress versus square root normalized displacements for TAMU footingsB-15
Figure B12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sandB-16
Figure B13 Summary of sand formation factors with corresponding CPT resistancesB-19
Figure B14 Normalized foundation stresses vs square root of normalized displacement for spread
footings on sand (modified after Mayne et al 2012) B-20
Figure B15 Definitions of capacity from three types of observed foundation load test responses
(modified after Kulhawy 2004) B-21
Figure B16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footingsB-22
B-1
Figure B17 Measured load vs displacement curves for 3 shallow foundations on clay at Baytown Texas
(Stuedlein amp Holtz 2010)B-25
Figure B18 Characteristic stress vs square root of normalized displacement for all three foundations at
Baytown site (Stuedlein amp Holtz 2010) B-26
Figure B19 Normalized foundation stress vs square root of normalized displacements B-27
Figure B20 Applied foundation stress vs CPT-calculated stress for 67 footingsB-28
Figure B21 Applied footing stress vs CPT calculated stress on logarithmic scales B-29
Figure B22 Summary graph and equations of direct CPT method for evaluating the vertical B-30
Figure B23 Foundation soil formation parameter hs versus CPT material index Ic B-32
Figure B24 Bearing capacity ratio qmaxqnet versus CPT material index IcB-33
Figure B25 Influence factor for rectangular foundations from elastic theory solution (Mayne amp
Dasenbrock 2017) B-34
List of Tables
Table B1 Direct CPT methods for bearing capacity of footings on clean sandsB-4
Table B2 Direct CPT methods for bearing capacity of footings on clays B-9
Table B3 Summary of large footings soil conditions and reference sources of database B-17
Table B4 Sources for shallow foundation performace B-34
B-2
B1 DIRECT CPT METHOD FOR SHALLOW FOUNDATIONS
B11 Introduction
The analysis of shallow foundations on soils is classically handled as a two-step and two-part set of
calculations involving (a) bearing capacity commonly reliant on limit plasticity solutions and (b)
settlement or more appropriately termed displacements that are assessed via elastic continuum theory
The utilization of cone penetration tests (CPT) provides the necessary geotechnical data for the site-
specific information on the subsurface conditions at the project site including the geostratigraphy and
evaluation of input parameters This is still a viable approach where the CPT readings are interpreted to
give the soil unit weight (γt) effective friction angle (ϕ) undrained shear strength (su) preconsolidation
stress (σp) and elastic moduli (D and E) for analysis
An alternative approach is the use of CPT results to provide direct assessments of bearing capacity
andor settlements A comparison of these two distinctly different and alternate paths is depicted in
Figure B1
Figure B1 Conventional methods versus direct CPT approach to shallow foundation response
A number of available direct CPT approaches for shallow footings are reviewed within this report
Moreover as the entire load-displacement-capacity response of shallow foundations to loading occurs
as a continuous nonlinear phenomenon a single direct CPT solution is presented for the behavior of
B-3
footings on sands (drained) and clays (undrained) The methods discussed herein are specifically to
address the general case of vertical loading of shallow foundations Additional more complex situations
that consider load eccentricity moments inclined forces sloping ground and other facets may be
handled using well-established procedures that are discussed elsewhere (eg Vesić 1975 Kulhawy et
al 1983)
This report focuses on foundations on granular soils andor soils exhibiting drained behavior since
Federal Highway Administration (FHWA) studies have concluded that less than 1 of shallow
foundations for highway bridges are placed on clay soils (Paikowsky et al 2010) One reason for this is
because soft-firm intact clays require a more complicated process that assesses both a short-term
analysis (undrained) as well as long-term analysis (drained) thus requiring undrained bearing capacity
undrained distortion displacements drained bearing capacity and drained primary consolidation
settlements
B12 Direct CPT Methods for Bearing Capacity of Footings on Sands
Classical bearing capacity solutions are based most often in limit plasticity theory albeit can be
formulated from limit equilibrium cavity expansion and numerical methods Because the limit
plasticity solutions are prevalent and adopt total stress analysis they are usually developed as either
fully drained (Δu = 0) or fully undrained (ΔVV = 0) where Δu equals excess porewater pressure and
ΔVV equals volumetric strain Paikowski et al (2010) provides an extensive review of various solutions
There are a number of direct CPT methods for the evaluation of the bearing capacity of footings and
shallow foundations Table B1 lists a variety of these direct CPT approaches for footings on sands In
the direct approach a one-step process is used to scale the measured cone penetrometer readings (ie
measured cone tip resistance (qc) total cone tip resistance (qt) sleeve friction (fs) andor measured
porewater pressure u2)) to obtain the ultimate bearing stress (qult) of the foundation
Table B1 Direct CPT methods for bearing capacity of footings on clean sands
Method Surface Footing Remarks and Embedment Notes
Meyerhof (1956) qult = qc (B12) middotcw qult = qc (B12)(1+DeB)cw
Note for silty sand reduce by 05
cw = 10 dry or moist sand cw = 05 submerged sand
Meyerhof (1974) qult = qc (B + D)40 with stresses in tsf
Presumably dry sands B = fdn width (feet) and D = depth (feet)
Schmertmann (1978)
NA (applies to embedded footings)
Square qult =055σatm(qcσatm)078
Strip qult =036σatm(qcσatm)078
See Figure A2
Embedment applies De gt 05(1+B) for Blt1m De gt 12 m for B gt1m
B-4
Canadian qult = Rk0middotqc Applied to FS = 3 Geotech Society where Rk0 = 03 where FS = factor of (CFEM 199 2) safety
Tand et al (1995) NA qult = Rkmiddotqc + σvo See Figure A3 where Rk = fctn(De B)
Frank and Magnan (1995)
qult = Rk0middotqc where Rk0 = fctn(sand
consistency) Loose Rk0 = 014
Medium Rk0 = 011 Dense Rk0 = 008
qult = Rk1middotqc + σvo where Rk1 = function (De B L amp
sand consistency) Factor Rk1 = Rk0[1+Rk206+04(BL) (DeB)]
and Rk2 = 035 (loose) 050 (medium) and 085 (dense)
Loose sand qc lt 5 MPa
Medium sand 8 MPa lt qc lt 15MPa
Dense sand qc gt 20 MPa
Eslaamizaad amp Robertson (1996)
qult = KΦmiddotqc See Figure A4
See surface footing equation KΦ = function (BDe shape and density)
Lee amp Salgado (2005)
qbL = βbcmiddotqc(AVG)
where qc averaged over distance B
Not addressed See Figure A5 for factor βbc = fctn(B DR
K0 and sB)
beneath the base
Eslami amp Gholami (2005
2006)
See embedded footing solution
qult = Rk1middotqc where Rk1 = function (ratio DeB
and normalized qcσvo) See Figure A6
Measured qc and qcσvo are geometric means over 2B deep
beneath footing
Robertson amp Cabal (2007)
qult = KΦmiddotqc with KΦ = 016
See surface footing equation
Briaud (2007) qult = KΦmiddotqc with KΦ = 023
Based on full-scale tests at Texas AampM
Mayne amp Illingworth
(2010) qult = 018 qc-Mean
Note qc -Mean is averaged CPT cone resistance over depth of
influence z = 15 B deep
Based on 30 footing load tests on 12
sands
Lehane (2013) qult = 016 qc-AVE Note qc-AVE is averaged CPT cone resistance within z (m) = [B (m)
]07
Based on 47 load tests
Table B1 Continued
Notes B = minimum footing width (or diameter) L = footing length De = embedment depth cw = water
table correction σvo = total overburden stress at bearing elevation and σvo = the effective vertical stress
B-5
Figure B2 Direct bearing capacity relationship for embedded square and strip footings
situated on clean sands (after Schmertmann 1978)
Figure B3 Direct CPT bearing factor Rk as function of footing width B and embedment depth
from finite element analyses (Tand et al 1995)
B-6
Figure B4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio
and sand consistency (Eslaamizaad amp Robertson 1996)
B-7
Figure B5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp
Salgado (2005) in terms of footing size relative density and base settlement
Figure B6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and
normalized cone tip resistance (after Eslami amp Gholami 2005)
B13 Direct CPT Methods for Bearing Capacity of Footings on Clays
For direct CPT evaluations of foundation bearing capacity in clays Table B2 summarizes some of the
well-documented methods that are available Generally these assume that the loading is fast enough
and the permeability of the clay is also sufficiently low such that an undrained (ie constant volume)
condition is maintained This is fine for short term loading of foundations particularly for soft to firm
intact clays However the undrained case is not permanent Given sufficient time excess porewater
pressures in the clay will eventually dissipate and hydrostatic conditions will return to equilibrium
During this dissipation phase primary consolidation will occur that results in drained settlements Also
a drained bearing capacity condition will prevail As the CPT is advanced at a constant rate of 20 mms
the recorded readings normally constitute undrained behavior from a direct measurement viewpoint
Thus direct CPT methods are normally focused at an undrained foundation response Nevertheless the
drained footing case can be addressed using CPT results via use of conventional analysis approach and
effective stress limit plasticity solutions that require piezocone penetrometers with porewater pressure
measurements
B-8
B
D)
L
B4060(3501320kc
Many of the earlier solutions were based on data from mechanical penetrometers andor older electric
cones that measure the uncorrected tip resistance (qc) It is now well-recognized that this measurement
must be updated to the corrected cone resistance (qt) because of porewater pressure effects that act on
the mechanics of the load cell and geometry of the particular penetrometer design The results are
particularly significant in clays silts and mixed soils that are soft to firm to stiff where excess porewater
pressures are recorded
Table B2 Direct CPT methods for bearing capacity of footings on clays
Method Footing Comments NotesRemarks
Meyerhof (1974) qult = αbcmiddotqc
where 025 le αbc le 050
Applicable to saturated insensitive clays under short-term loading
Cone tip resistance
measured by
mechanical CPT
Trofimenkov (1974)
Approximation
qult asymp (qc33)09
(Assuming FS = 3)
Strip footings on clays and sandy clays 06m le B le 15m 1m le zemb le 25m
Mechanical CPT
with qc and qult in
kgcm2
Schmertmann (1978)
qult = function (qc and foundation shape)
Relationship shown in Figure A7 for square and strip footings
Cone tip resistance
measured by
mechanical CPT
Tand et al (1986)
qult = Rkmiddot(qc - σvo) + σvo
Factor Rk obtained from Figure A8 accounts for surface and deep footings Separate Rk factor for intact and jointed clays
= (qc1middotqc2)05 qc
where qc1 is geometric mean from bearing elevation to 05B deeper and qc2 is geometric mean from 05B to 15B beneath foundation base
Mix of data from
mechanical CPT qc
and electric CPT qc
LCPC Method
(Frank amp Magnan 1995)
Footing a surface
qult = kc middotqc + σvo
where kc = 032
Embedment case
Note Methods above generally consider undrained bearing capacity
B-9
Figure B7 Direct relationship between ultimate bearing stress in clays and measured cone tip
resistance (Schmertmann 1978)
B-10
Figure B8 Direct relationship between ultimate foundation bearing stress in clays and net
cone tip resistance (after Tand et al 1986)
B14 Direct CPT Assessments of Settlement and Displacements of Shallow Foundations
Methods for evaluating the magnitude of foundation displacements (s) can be found using
elastic theory subgrade reaction methods spring models and numerical simulations The
classical approach to footing settlement calculations on sands is to utilize elastic theory
solutions (eg Poulos amp Davis 1974) which take on the form
119902∙119861∙119868∙(1minus1205842) 119904 = B1
119864119904
where q = applied footing stress and I = displacement influence factor from elasticity theory
The value of I depends upon foundation geometry footing rigidity embedment depth variation of soil
modulus with depth compressible layer thickness and other factors as discussed by Mayne amp Poulos
(1999) For instance for a flexible circular footing of diameter B resting on an infinitely deep soil
formation having a constant modulus Es with depth the influence factor I is equal to 1 for the center
point The results of in-situ field tests such as pressuremeter tests (PMT) standard penetration tests
(SPT) cone penetration tests (CPT) flat dilatometer tests (DMT) and other methods can be used to
ascertain the input value of soil modulus Es
Alternatively direct CPT methods have been investigated to provide a one-step assessment of
foundation displacements Selected methods for direct CPT evaluation of shallow foundation
settlements on granular soils is include DeBeers amp Martens (1957) Meyerhof (1965) DeBeer (1965)
Thomas (1968) Schmertmann (1970) Meyerhof (1974) Berardi amp Jamiolkowski amp Lancellotta (1991)
Robertson (1991) Lutenegger amp DeGroot (1995) Lehane amp Dougherty amp Schneider (2008) Gavin amp
Adekunte amp OrsquoKelly (2009) Mayne amp Illingworth (2010) and Uzielli amp Mayne (2012)
B141 Large Scale Footing Load Tests
The measured load-displacement response of shallow foundations is conducted in the field using either
stepped loading procedures or continuous rate of displacement methods Results from the well-
documented test program at the Texas AampM University (TAMU) site (Briaud amp Gibbens 1999 Briaud
2007) for five spread footings on sand are shown in Figure B9 Each footing clearly shows a nonlinear
B-11
behavior to loading over the range of testing This site is one of the national geotechnical
experimentation sites (NGES) funded by the National Science Foundation (NSF) FHWA and American
Society of Civil Engineering (ASCE)
Figure B9 Measured load-displacements for each of the five large footings at TAMU (Briaud
amp Gibbens 1994) sand site
B142 Generalized Direct CPT Method for Footing Response on Soils
The use of applied foundation stress versus normalized displacement curves (q vs sB) is generalized to
footings on sands silts intact clays and fissured clays A square root plotting of the normalized
displacements (sB) permits a single parameter characterization for each specific soil type that in turn
correlates with the net cone tip resistance The generalization is based on a statistical review of data
from large foundations (B gt 05m) involving 70 full-scale load tests with available CPT data For fine-
grained soils the data are from electronic piezocone tests where the proper correction from qc to qt has
been obtained to allow the best possible results The methodology permits an evaluation of the load-
displacement-capacity response of shallow square footings based on CPTs performed in the four types
B-12
of ground conditions For the TAMU footings the characteristic stress-normalized displacement
response is shown in Figure B10 where all five foundations can be represented by a single curve
Figure B10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU
(Briaud amp Gibbens 1994) footings
B143 Characteristic Stress-Displacement Curves
Fellenius amp Altaee (1994) recommended the concept of a characteristic stress (q) vs normalized
displacement (sB) curve for foundation response on a given soil formation later supported by Decourt
(1999) Briaud amp Gibbens (1999) Lutenegger amp Adams (2003) and Briaud (2007) In this approach the
measured load (Q) vs displacement from individual footings of various sizes that rest on the same soil
conditions collapse to a unified relationship given in Equation B2
119939119943 119956 119954119938119953119953119949119946119942119941 = 119938119943 (119913
) B2
B-13
where af and bf are empirical fitting coefficients (Decourt 1999 Uzielli amp Mayne 2011 2012)
In a review of measured load tests data from large spread footings situated on different sands it has
been suggested that Equation B2 can be reduced to a single parameter expression by use of square root
plotting where the exponent bf is equal to 05 (Mayne amp Illingworth 2010 Mayne et al 2012)
119956 119954119938119953119953119949119946119942119941 = 119955119956radic
119913 B3
where rs = fitted soil parameter from best fit line regression In the case of sands the coefficient rs
represents the supporting soil conditions including particle size relative density and fines content as
well as geologic origin aging and overconsolidation effects
The results from the five TAMU footings are presented in this format in Figure B11 that shows the
formation factor rs = 486 MPa for this sand deposit This single coefficient can be used to express the
nonlinear load versus settlement of any size footing on this sand site Moreover one criterion from
Eurocode for bearing capacity that is used by the European community identifies that stress (or load)
which corresponds to (sB) = 10 That is when the displacement equals ten percent of the footing
width then the capacity has been reached Therefore adopting this criterion for footings on sands
the ultimate bearing stress (qult) for footings on sand can be taken when (sB) equals 010 or when
square root of (sB) equals 0316 In that case this gives qult equal to 0316 middot rs For the TAMU site this
gives qult equal to 0316 times (486) which equals 153 MPa as illustrated by Figure B12
B-14
Figure B11 Characteristic stress versus square root normalized displacements for TAMU
footings
B-15
Figure B12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sand
B15 Footing Database
A database of spread footings and large plates was assembled which considers only full-size shallow
foundations (05 le B le 6 m) that rest on sands silts and clays Sites for the foundations were also subjected to CPT The inclusion solely of large footings are significant because results from small-scale
model footings exhibit scale effects In fact experimental centrifuge work and numerical simulation
studies have shown that bearing capacity factors are size-dependent especially for factor Nγ decreasing
with footing width B (Kimura et al 1985 Cerato amp Lutenegger 2007 Mase amp Hashiguchi 2009) A
direct approach based on full-size footings provides a more reliable means for foundation evaluation
The compiled database is given in Table B3 with a total of 70 large footings and plates These include a
listing of 34 foundations on 13 sands 11 footings on 4 silt deposits 13 footings or large plate load tests
on 6 intact clays and 12 foundations on 5 fissured clays Most of the footings were square or nearly
square (80) while the remaining were circular The largest foundation was a mat 5 m by 14 m in plan
loaded by stacking Kentledge concrete blocks to cause a bearing failure in the underlying soft clay For
sands the largest footing consisted of the 6-m square concrete pad In terms of equivalent square
footings mean footing sizes were 155 m (sands) 114 m (silts) and 126 m (clays)
B-16
Additional details on the individual load tests and site conditions are given elsewhere (Mayne 2009
Mayne amp Illingworth 2010 Uzielli amp Mayne 2011 2012 Mayne et al 2012 Mayne amp Woeller 2014)
Table B3 Summary of large footings soil conditions and reference sources of database
Sand Site Location Soil Conditions Footings Numbers
Shapes and Sizes
ReferencesSource
College Station
Texas Pleistocene sand 5 Square 1 15 25 33 m Briaud amp Gibbens (1999)
Kolbyttemon Sweden Glaciofluvial sand 4 Rect B = 06 12 17 24 m
Bergdahl et al (1985 1986)
Fittija Sweden Glaciofluvial sand 3 Rect B = 06 17 24 m Bergdahl et al (1984 1985)
Alvin West Texas Alluvial sand 2 Circular D = 235 m Tand et al (1994)
Alvin East Texas Alluvial sand 2 Circular D = 22 m Tand et al (1994)
Perth Australia Silceous dune sand 4 Square B = 05 and 10 m
Lehane (2008)
Grabo T1C Sweden Compacted sand fill
1 Square B = 046 m Long (1993)
Grabo T2C Sweden Compacted sand fill
1 Square B = 063 m Long (1993)
Grabo T3C Sweden Compacted sand fill
1 Square B = 080 m Long (1993)
Labenne France Dune sand 6 Square B = 07 and 10 m
Amar et al (1998)
Green Cove Florida Brown silty sand 1 Circular D = 182 m Anderson et al (2006)
Durbin South Africa
White fine sand 1 Square B = 609 m Kantley (1965)
Porto Portugal Residual clayey sands
1 Circular D = 12 m and 1 plate
Viana da Fonseca (2003)
Jossigny France Soft clayey silt 2 Square B = 1 m Amar et al (1998)
Tornhill Sweden Glacial Baltic till 3 Square B = 05 1 2 m Larsson (2001)
Vagverkel Sweden Stiff medium silt 3 Square B = 05 1 2 m Larsson (1997)
Vattahammar Sweden Brn-Gry layered silt 3 Square B = 05 1 2 m Larsson (1997)
B-17
Table B3 Continued
Clay Site Location Soil Conditions Footings Numbers Shapes and Sizes
ReferencesSource
Baytown Texas Fissured Beaumont 2 plates with d = 076 m Stuedlein amp Holtz (2008)
Belfast Ireland Soft clay silt ldquosleechrdquo
1 Square Pad B = 20 m Lehane (Geot Engr 2003)
Bothkennar Scotland Soft silty clay 2 Square Pads B = 22 24 m
Jardine et al (1995)
Bangkok Thailand Soft to stiff clay 4 Square 067 075 090 10 m
Brand et al (1972)
Haga Norway Stiff OC clay 2 Square Footings B = 10 m
Andersen amp Stenhammar (1982)
Rio Grande Brazil Sandy residual clay 3 Square 04 07 and 10 m
Consoli et al (1998)
Shellhaven England Soft estuarine clay Rectangular 5 m by 14 m Schnaid et al (1993)
Texas City A Texas Coast Fissured Beaumont 3 circular plates d = 058 m
Tand et al (1986)
Texas City B1 Texas Coast Fissured Beaumont 2 circular plates d = 058 m
Tand et al (1986)
Texas City B2 Texas Coast Fissured Beaumont 1 circular plate d = 058 m
Tand et al (1986)
Alvin Texas Texas Coast Fissured Beaumont 3 circular plates d = 058 m
Tand et al (1986)
For the series of footings on sands Figure B13 shows the summary of measure formation factors (rs) for
the 34 foundation load tests Also indicated on the graph are the corresponding mean qc values of the
sands averaged over 15middotB beneath the bearing elevations
Note also in sands that the qt and net resistance (qnet) are all very close to the measured qc since (a)
porewater pressure corrections for the a-net value are negligible in clean sands and (b) overburden
stresses are typically only 1 to 3 of the magnitude of qc This is especially true of shallow depths
Therefore for clean sands qnet is approximately qt which is approximately qc
B-18
Figure B13 Summary of sand formation factors with corresponding CPT resistances
As suggested by Briaud (2007) various in-situ measurements can be used to normalize the results from
cone penetration tests in this case Herein the qc from the CPT soundings in the sands are used to
normalize the footing stress axis as presented in Figure B14 It is evident that the results of the load-
displacement response of all 34 footings on 13 different sands can be captured by the characteristic
stress versus normalized displacement with the inclusion of the site-specific cone tip resistance In this
case the mean trend for all footings and sand sites indicates a simple expression shown in Equation B4
119954119943119952119952119957119946119951119944 119956 Clean quartz-silica sands = 120782 120787120790120787radic B4
119954119940 119913
B-19
Figure B14 Normalized foundation stresses vs square root of normalized displacement for
spread footings on sand (modified after Mayne et al 2012)
The results of measured force-displacement curves (Q vs s) from footing load tests are nonlinear and
thus the definition of capacity must be addressed Kulhawy (2004) identified three main curve types
that occur during load testing depicted in Figure B15 The most common response (Type A) shows no
clear peak and is observed in sands and silts as well as slow loading of insensitive clays and fissured
clays In this case ldquocapacityrdquo can be defined by the Eurocode corresponding to the load (or stress) when
sB=10 (Amar et al 1998) For foundations situated on many clays a plateau capacity is reached as
depicted as Type B response In rare cases where sensitive clays or structured soils are encountered a
Type C relationship occurs whereby the load-displacement curve reaches a peak followed by strain
softening In the one case observed in this study for Type C loading (Haga clay) the corresponding
ldquopeak capacityrdquo was reached at a pseudo-strain of sB = 4 as shown in Figure B16 followed by some
strain softening In the cases of soft clays at Belfast and Bothkennar a plunging type (plateau failure)
was achieved at around sB=7
B-20
Figure B15 Definitions of capacity from three types of observed foundation load test
responses (modified after Kulhawy 2004)
B-21
Figure B16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footings
In the case of fine-grained soils that develop excess porewater pressure during cone penetration the
use of the net total cone resistance (qtnet) which equals the cone tip resistance minus the total stress
must be considered (Lunne et al 1997) As noted earlier the correction of CPT data in sands is minimal
because of the low value of penetration porewater pressures and the fact that total overburden stress is
small relative to the cone resistance especially for shallow soundings Thus the use of qtnet may be
substituted for qc into the trend of Figure B1
B16 Undrained and Drained Capacity
The footings on sands typically show drained response with no excess porewater pressures developed
during loading For the foundations situated on silts analyses by Larsson (1997 2001) concluded that
these too show drained conditions For shallow foundations on clays the entire range of drainage cases
are possible ranging from undrained to partially-drained to fully-drained depending upon the applied
rate of loading relative to the permeability of the geomaterial If sufficient data were available for each
of the clay sites then the recent evaluation of drainage condition could be assessed using normalized
velocity criteria (Chung et al 2006) However this was not possible with the current data set Instead
B-22
the test loadings of footings on clays have essentially been assumed to primarily occur under undrained
conditions following recommendations by Tand et al (1986) For this case a limiting bearing stress can
be calculated from bearing capacity analyses and related directly to the CPT readings
From classical bearing capacity calculations using limit plasticity solutions the ultimate foundation stress
is shown in Equation B5
= 119925119940 ∙ 119956119958 B5 119954119958119949119957
where Nc equals the bearing factor for constant volume (514 for strip and 614 for square or circular
plan) and su equals the undrained shear strength of the clay For the case of soft to firm clays of low to
medium sensitivity the operational strength can be related to the preconsolidation stress (Mesri 1975
Jamiolkowski et al 1985) as shown in Equation B6
prime 119956119958 = 120782 120784120784120648119953 B6
Finally the effective preconsolidation stress has been linked directly to the net cone resistance (Chen amp
Mayne 1996 Demers amp Leroueil 2002) as shown in Equation B7
prime 120648119953 = 120782 120785120785(119954119957 minus 120648119959119952) B7
Combining Equations (B5) (B6) and (B7) results in direct expressions for undrained bearing capacity on
intact clays from CPT results as shown in Equations B8 and B8-2
Strip footing 119954119958119949119957 = 120782 120785120789120785(119954119957 minus 120648119959119952) B8
B-23
Squarecircle 119902119906119897119905 = 0445(119902119905 minus 120590119907119900) B8-2
Equation (B8-2) is in excellent agreement with Tand et al (1986) database study for the case of intact
clays and footings with no embedment Their study also provided a lower relationship for foundations
situated on jointed clays specifically recommending that the bearing stress not exceed 030 qtnet
Review of the available data in Figure 3 shows this to be rather conservative and a more realistic value
may be taken at 040 qtnet for fissured clays
B161 Normalized Undrained Footing Response
The characteristic stress vs normalized displacement curves for footings on clays exhibiting undrained
conditions can be verified using a relatively recent case study on Beaumont clay by Stuedlein amp Holtz
(2010) The Baytown Texas site was investigated using an exploration program of soil borings
piezocone soundings and laboratory testing The field testing included a large square footing (B = 276
m) and two small circular plates (D = 076 m) The measured load-displacement responses for all three
foundations are presented in Figure B17 Of course the large footing carried considerably higher loads
than the two small plates on the same soil deposit Yet using the aforementioned characteristic stress
(q) vs square root of normalized displacement (sB) essentially gave a unique relationship for all 3
foundations as presented in Figure B18 In this case utilizing the form of Equation B2 the
characteristic coefficient rs is 177 MPa for this deltaic clay formation
B-24
Figure B17 Measured load vs displacement curves for 3 shallow foundations on clay at
Baytown Texas (Stuedlein amp Holtz 2010)
B-25
Figure B18 Characteristic stress vs square root of normalized displacement for all three
foundations at Baytown site (Stuedlein amp Holtz 2010)
B17 General Direct CPT Method
In a manner analogous to the normalization of applied foundation stresses shown in Figure B14 for
footings on sands a generalized direct CPT method for shallow foundations on different soils can be
made in Equation B9
120782120787 119956 119954 = 119945119956 ∙ 119954119957119951119942119957 ∙ (119913
) lt 119954119940119938119953119938119940119946119957119962 B9
where qcapacity is equal to the foundation bearing capacity of the ground and hs is equal to the empirical
fitting term that depends on soil type Specifically hs equals 058 for sands 112 for silts 147 for
fissured fine-grained soils and 270 for clays A summary of the normalized stress versus the normalized
displacement curves for these four soil types is presented in Figure B19 where the data fitting to obtain
the hs parameter on clays are limited to (sB) lt 004 corresponding to undrained loading The data on
B-26
silts and sands are considered fully drained whereas the fissured clay subset may be partially drained to
undrained The statistical measures for obtaining the fitted parameter hs are quite good as shown in
Figure B20 with the coefficient of determinations (r2) values of 0947 for sands 088 for silts 0935 for
fissured and 0925 for intact clays Additional statistical evaluations on the database are provided by
Uzielli amp Mayne (2011)
Figure B19 Normalized foundation stress vs square root of normalized displacements
B-27
Figure B20 Applied foundation stress vs CPT-calculated stress for 67 footings
The same data are shown on Figure B21 in a log-log plot to emphasize the early parts of the load-
displacement responses of these footings The best fit line from regression analyses shows excellent
statistical correlations As detailed earlier the undrained bearing capacity of square or circular footings
on clays can be taken as approximately 040 times qtnet For silts and sands that experience drained
loading with no excess porewater pressures being developed and no clear peak value for capacity the
(sB) =10 criterion can be used In this case Equation 7 can be used directly to obtain stress q equal to
qcapacity when (sB)05 is 0316
An alternate approach is presented in Figure B22 summarizing the direct CPT method for estimating the
load-displacement-capacity of shallow square footings on four soil types The maximum values of soil
bearing capacity (qmax) can be capped at stresses corresponding to a percentage of the average
measured qtnet determined over the depth interval from the foundation bearing elevation to 15middotB
below the foundation In such an approach the foundation bearing capacity can be taken as
(qcapacityqtnet) of 02 for sands 035 for silts 040 for fissured and 045 for intact clays Using the assigned
values of the hs parameter the corresponding cut-off for the general stress-normalized displacement
relationship given by Equation 7 can be established specifically giving corresponding values of the
B-28
normalized displacements which can also be considered as pseudo-strains equal to (sB) max of 4 for
clays 7 for fissured clays 10 for silts and 12 for sands
Figure B21 Applied footing stress vs CPT calculated stress on logarithmic scales
B171 CPT Soil Material Index Ic
The soil behavioral type can be assessed indirectly by CPT using a material index (Ic) as defined by
Robertson (2009a b) shown in Equation B10
119920119940 = radic[120785 120786120789 minus 119949119952119944119928119957119951]120784 + [120783 120784120784 + 119949119952119944119917119955]120784 B10
B-29
Figure B22 Summary graph and equations of direct CPT method for evaluating the vertical
where Qtn equals the stress-normalized cone tip resistance and Fr equals the normalized sleeve friction
calculated from the cone penetrometer readings as shown in Equations B11 and B12 respectably
(119954119957minus120648119959119952)120648119938119957119950 = 119951 B11 119928119957119951 prime (120648119959119952120648119938119957119950)
119943119956 119917119955() = 120783120782120782 ∙ B12 (119954119957minus120648119959119952)
where σatm equals a reference stress equal to atmospheric pressure (σatm = 1 atm asymp 1 bar asymp 100 kPa) In
the initial evaluation the exponent n is set to n = 1 to find the soil behavioral type (SBT) based on a 9-
zonal chart as shown in Figure 23 The value of exponent n varies with soil type ranging from n = 1 in
intact clays and decreasing with increasing grain size to around approximately 075 in silts and
B-30
approximately 05 plusmn 02 in clean sands The appropriate value of exponent n is found by iteration until
conversion using the relationship (Robertson 2009b) shown in Equation B13
prime 120648119959119952 119951 = 120782 120785120790120783 ∙ 119920119940 + 120782 120782120787 ( ) minus 120782 120783120787 le 120783 120782 B13 120648119938119957119950
As the distinction that footings on intact clays subjected to fast loading are behaving primarily under
undrained conditions (ie constant volume) Robertson (2009a) suggested that this occurs when Ic gt 27
while in contrast Ic lt 25 corresponds more or less to drained loading (ie no excess porewater
pressures)
The CPT material index can be used to identify soil type and the corresponding characteristic hs value for
estimating footing load-displacement response In a number of the case studies investigated the results
of the sleeve friction readings were also available to allow the calculation of Ic with depth at these sites
Figure B23 shows the tentative relationship between the values of the hs parameter plotted versus the
corresponding CPT material index with an approximate trend given by Equation B14
120784120785 119945119956 = 120784 120790 minus 120783120787 B14
119920119940 120783+( ) 120784120786
B-31
Figure B23 Foundation soil formation parameter hs versus CPT material index Ic
Soil behavioral type from the Ic ranges given by Robertson (2009b) and are shown in Figure B23 with an
overall general agreement with the known soil classifications In summary the CPT can provide an
evaluation of the load-displacement-capacity response directly via Equation B15 which may be of
interest in mixed soil types such as silty sands sandy silts and the like
120784120785 Footing stress 119954 = 119954119951119942119957 ∙ radic(119956frasl119913) ∙ [120784 120790 minus
)120783120787] B15
120783+(119920119940frasl120784120786
As discussed earlier foundation bearing capacity may be taken by a limiting value of pseudo-strain (sB)
max or by qmax as specified as a percentage of qtnet This definition of bearing capacity can be seen in
comparison to soil type as seen in Figure B24
B-32
Figure B24 Bearing capacity ratio qmaxqnet versus CPT material index Ic
B172 Rectangular Foundations
This same elastic solution from Giroud (1968) for a CPT method for square and circular foundations were
utilized for rectangular foundations The study covered rectangular distortions (AB) ranging from 1
(square) to very long foundations with AB equal to 20 where A represents the foundation length and B
represents the foundation width (Mayne amp Dasenbrock 2017) The influence factor for rectangular
shaped foundations is given in Figure B25 and the expression in Equation B16
= (119912119913)120782120785120786120787 119920119912119913 B16
B-33
Figure B25 Influence factor for rectangular foundations from elastic theory solution (Mayne
amp Dasenbrock 2017)
Including 32 full scale load tests on sands results from an additional 98 shallow foundations have been
reported in previous studies The foundations were primarily square or circular and among these include
large spread footings with measured settlements and bearing capacity Table B4 shows a summary of
the number of footings and footing sizes used in the study The range of length to width ratios varied
from 1 lt (AB) lt 23 with an average value of (AB) equal to 238 The embedment depth to footing
width ranged from 0 lt DeB lt 222 and averaged 046
Table B4 Sources for shallow foundation performace
Source of Data No of
Footings
Mean B
(m)
Max B
(m)
Min B
(m)
Mean A
(m)
Max A
(m)
Min A
(m)
Mayne et al (2012) 32 149 609 046 149 609 046
Lehane (2011) 8 041 027 060 041 060 027
Gifford et al (1987) 17 846 1588 527 1591 3596 701
Jeyapalan amp Boehm
(1986)
13 1022 2892 400 1671 3049 640
B-34
Schmertmann
(1970)
31 945 5608 061 1288 8668 061
Papadopoulos
(1992)
29 870 3600 100 1300 7290 100
Total 130 671 5608 027 1012 8668 027
The solution for shallow foundation settlements given back in Equation B1 combined with
the expression in Equation B16 takes on the form
119954∙119913∙119920119912119913∙(120783minus120642120784) 119956 = B17
119916119956
Combining the direct CPT equation developed from the 32 full scale load tests (Equation B15)
together with the influence factor for rectangular shaped foundations becomes
minus120782120785120786120787 120784120785 119912 Footing stress 119954 = 119954119951119942119957 ∙ radic(119956frasl119913) ∙ [120784 120790 minus
)120783120787] ∙ [ ] B18
120783+(119920119940frasl120784120786 119913
B18 Conclusions
A direct CPT method for square rectangular and circular shallow footings is developed using a database
of 166 full-scale field load tests where the minimum footing size of an equivalent square width of 05 m
is established to avoid scaling issues The use of load vs displacement is generalized by characteristic
stress vs normalized displacement (sB) and further simplified by a square root plotting technique that
captures the ground response in a single parameter that is related to the qtnet Data are grouped
according to four main soil categories sands silts fissured clays and intact clays It is generally believed
that the footings on sands and silts exhibit fully drained behavior while in intact clays an undrained
response occurs under conditions of constant volume
B-35
The summary equation for evaluating the vertical stress-displacement-capacity of square rectangular
and circular footings is given by
minus120782120785120786120787 119956 119912 119954119943119952119952119957119946119951119944 = 119945119956 ∙ 119954119957119951119942119957 ∙ radic ∙ ( ) lt 119954119950119938119961 B19
119913 119913
where the parameter hs also relates directly to Ic The foundation capacity depends upon the mode of
failure as well as drainage conditions and can be taken as a fraction of qtnet where qmaxqtnet is 020 in
sands 035 in silts 040 in fissured clays and 045 in intact clays as shown in Figure 24 Alternatively the
capacity can be defined using a limiting pseudo-strain given by (sB) max of 4 in clays 7 in fissured
10 in silts and 12 in sands
B-36
APPENDIX C
List of Figures
Figure C1 Components of Axial Pile Capacity C-3
Figure C2 Selected alpha relationships as function of the undrained shear strengthC-5
Figure C3 Pile side friction coefficient β in terms of effective friction angle and overconsolidation ratio
C-11
Figure C4 Concept of Direct versus Rational Method for CPT evaluation of axial pile capacityC-14
Figure C5 Soil behavior type using CPT via original UniConeC-19
Figure C6 Soil behavior type using CPT via Modified UniCone C-19
Figure C7 Nine-zone soil behavioral soil type using normalized piezocone parameters C-21
Figure C8 Summary of Modified UniCone Method for direct CPT assessment of axial pile capacity C-22
Figure C9 Elastic continuum solution for axial pile response under compression and tension loadingC-24
List of Tables
Table C1 Alpha Methods for Axial Pile Side Friction where fp = αmiddotsu C-5
Table C2 Beta Methods for Axial Pile Side Friction where fp = βmiddotσvoC-9
Table C3 Selection of Direct CPT Methods for Axial Pile CapacityC-15
C-1
C1 EVALUATING DEEP FOUNDATION RESPONSE FROM CONE
PENETRATION TESTS
C11 Introduction
The axial response of deep foundations includes the load-displacement-capacity and axial load transfer
when driven pilings and drilled shafts are loaded in compression and uplift The use of cone penetration
testing (CPT) especially seismic piezocone tests (SCPTu) are advantageous since they provide
information on the subsurface soils and their geomaterial properties The SCPTu provides at least four
measurements on soil behavior with depth including (a) cone tip resistance qt (b) sleeve friction fs (c)
penetration porewater pressure u2 and (d) shear wave velocity Vs This offers the opportunity to
determine the geostratigraphy unit weight effective overburden stress shear strength stress state
and stiffness of the ground from a single exploratory sounding thus values in economic expediency
and reliability The axial pile capacity can be calculated based on static equilibrium of forces acting along
the sides and base of the pile foundation The displacement of the pile can be ascertained using
elasticity theory from closed-form solutions boundary elements andor finite element analyses Load
transfer occurs along the length of the pile and only a portion of axial forces are transmitted to the base
or toe or tip of the pile These too can be assessed using elasticity solutions
An alternate approach to pile capacity involves the utilization of direct CPT methods available
since the 1970s but in the past 10+ years a number of new and statistically reliable algorithms
have been developed which can be implemented for highway design and construction
C111 Axial Pile Capacity
The axial compression capacity of a single pile foundation is composed of a shaft or side component and
end-bearing component at the base as depicted in Figure 1 For a circular pile the side capacity (Qs) is
determined from the unit side friction (fp) acting along the surface area of the shaft which is As = πmiddotdmiddotL
where d = pile diameter and L = length embedded below grade
If the magnitude of side friction is uniform and constant with depth the side capacity is simply
C-2
119928119956 = 119943119953 middot 119912119956 C1
Moreover however many piles extend through multiple layers and a summation of unit side
frictions acting on various pile segments must be tabulated over the length of the pile as
suggested by Figure C1
Figure C1 Components of Axial Pile Capacity
The unit-end bearing resistance (qb) acts over the base of the pile tip where the area of a circular pile is
given by Ab = πmiddotd24 For piles in compression loading the base capacity is determined from Equation
C2
119876119887 = 119902119887 middot 119860119887 C2
and for piles in tension (or uplift) it is normally taken that Qb = 0
C-3
C112 Pile Unit Side Friction
Several different approaches can be adopted for evaluating the unit pile side friction (fp) prior to full-
scale load testing and construction (Poulos amp Davis 1980 ONeill 2001) The most common types
including the alpha and beta methods as summarized in Tables 1 and 2 respectively In the alpha
method an empirical coefficient (α) is applied to the undrained shear strength (su) of clay soils to
obtain
119891119901 = 120572 middot 119904119906 C3
An illustrative example of alpha curves is shown in Figure C2 A difficulty with the alpha
method is the evolution of many variants and changes to the expressions for its estimation It is
also restricted in its use for specific types of piles in clays and fine-grained soils
C-4
Figure C2 Selected alpha relationships as function of the undrained shear strength
Table C1 Alpha Methods for Axial Pile Side Friction where fp = αmiddotsu
C-5
Reference
Source
Alpha Equation Remarks
Tomlinson
(1957)
2716801102
uu ssAll pile types (steel
concrete timber)
Note su in ksf
Tomlinson
(1957)
2616201102
uu ssConcrete piles
Note su in ksf
McClelland
(1974)
Kerisel α = 07 - 031middotln(su)
Peck α = 10 - 02middotsu
Woodward α = 091middot(su)091
Driven piles in clay
Note su in ksf
Semple
(1980) )1(511
)1(1
OCR
OCR Summary from 9 series of
pile load tests
American
Petroleum
Institute (API
1981)
For suσvo le 035 α = 10
For suσvo gt 080 α = 05
Otherwise α = 1 + 1111middot (035 - suσvo)
Driven steel pipe piles
Tomlinson
(1986)
1 α = 55 (su)-091
2 α = 785 (su)-102
3 α = 605 (su)-100
where 75 lt su (kPa) lt 210
1 Driven piles
2 Bored piles
3 Driven piles in till with L
lt 10middotd
API (1987) 1 α = 10 for su lt 25 kPa Driven piles in clay other
than Gulf of Mexico
C-6
2 α = 125 - 001middotsu for 25 lt su lt 75 kPa
3 α = 050 for su gt 75 kPa
Kulhawy amp
Jackson
(1989)
α = 021+026middot(σatm su) le 1 Analyses of 106 drilled
shaft foundations
Table C1 Continued
API (1989) 05 For suσ vo le 1 α =
su radic frasl prime σvo
minus025 su For suσ vo gt 1 α = 05 ∙ ( frasl prime ) σvo
Driven steel pipe piles
Kulhawy amp
Jackson
(1989)
OCR
OCR
sin50
tan)sin1( sin
Λ = (1-CsCc) asymp 080 for
insensitive clays
Nowacki et al
(1996)
minus05 su For suσvo le 07 α = 05 ∙ ( frasl prime ) σvo
minus02 su For suσvo gt 07 α = 055 ∙ ( frasl prime ) σvo
Driven pile foundations
Kolk and van
der Velde
(1996)
α =09middot FL middot (suσvo)-03 lt 10
FL = [ (L - z)d]-02 = length term
L = pile length
Note su obtained from
UU lab tests
Applicable to driven piles
in clays
C-7
Knappett amp α = 1 for su le 30 kPa Non-displacement piles in
Craig (2012) fine-grained soils α = 116 - (su185) for 30 kPa le su le 150 kPa
α = 035 for su gt 150 kPa
Knappett amp
Craig (2012) α = 055 ∙ 02 40
( ) (LfraslD)
∙ su ( frasl prime σvo
minus03
) Displacement piles in fine-
grained soils
z = depth at point considered
d = outside diameter of pile
Karlsrud et al
(2005 NGI
Method)
α = function (suσvo and plasticity index) Driven pilings
Fleming et al
(2009) 05 minus05 su su For su σvo le 1 120572 = ( frasl prime ) ∙ ( frasl prime )
σvo NC σvo
05 minus025 su su For su σvo gt 1 α = ( frasl prime ) ∙ ( frasl prime ) σvo NC σvo
For driven piles in clay
where (suσvo)NC is the
normalized shear strength
for normally-consolidated
clay
Brown et al
(2010)
α = 0 for 0 lt z le 5 feet
α = 055 for z gt 5 feet and (suσatm) le 15
α = 055-01middot(suσatm -15) for 15le (suσatm) le 25
Drilled Shafts and Bored
Piles
Table C1 Continued
C-8
OCRK )sin1(0
Karlsrud
(2012)
α = function (su σvo and plasticity index) Driven pilings
Note as reported by Karlsrud (2012)
In the beta method the coefficient β is applied to the effective overburden stress (σvo) at the point of
concern along the pile length
prime = 120573 C4 119891119901 middot 120590119907119900
Table C2 Beta Methods for Axial Pile Side Friction where fp = βmiddotσvo
Reference Source Beta Equation Remarks
Burland (1973) β = (1-sinϕ) middot tan ϕ Piles in NC clays
Meyerhof (1976) β = Kmiddottanδ where K = K0 = for NC clays
K = 15middotK0 for OC clays
Poulos amp Davis
(1980)
β = (1-sinϕ) middot tan ϕ middot OCR05 Piles in OC stiff clays
Table C2 Continued
C-9
Kulhawy amp Jackson
(1989)
β = (1-sinϕ) middot tan ϕ middot OCRsinϕ Piles in quartz sands and insensitive
clays
ONeill (2001) β = (1-sinϕ) middot tan θ middot OCRsinϕ Drilled shafts where θ = (δϕ)middotϕ and
δ = interface friction between pile and soil
Fleming et al
(2009)
β = Kmiddottanδ K = lateral stress coefficient and δ =
friction angle between soil and pile
material
Karlsrud (2012) β = fctn (OCR and PI) PI = plasticity index of the clay ()
Mayne and Niazi
(2017)
β = CM middot CK middot K0 middot tanϕ
where K0 = (1-sinϕ) middot OCRsinϕ for soils
that are virgin loaded then unloaded
CM = pile material factor = 1 (drilled
augered) 09 (prestressed or precast
concrete) 08 (timber) and 07
(steel) and CK = pile installation factor
= 09 (bored or augered) 10 (low
displacement eg H-pile or open-end
pipe) and 11 (driven high
displacement eg prestressed
concrete closed-end pipe)
As the beta approach applies to all types of soils (gravels sands silts and clays) and more or
less has remained unchanged since its advent circa 1970 it has been selected for further
discussion herein In the direct form for calculation of unit pile side friction the expression is
given by
119874119862119877119904119894119899120601 prime prime 119891119901 = 119862119872 ∙ 119862119870 ∙ (1 minus 119904119894119899120601prime) middot ∙ 119905119886119899120601prime ∙ 120590119907119900 C5
C-10
where CM is a soil-pile interface friction coefficient and CK = pile installation factor Values of CM are in
the range 07 lt CM lt 10 depending upon pile material (steel wood concrete) and ranges for CK vary
09 lt CK lt 11 depending upon method of installation (auger drill driven) as detailed in Table 2
The value of K0 is limited to the passive stress coefficient which for the simple Rankine case is given by
KP = (1+sin ϕ) (1 - sin ϕ ) Thus there is a maximum value of overconsolidation ratio for which
Equation C5 applies given by OCRlimit = [(1+sin ϕ ) (1 - sin ϕ )2] (1sinϕ)
The corresponding graph in Figure C3 shows the relationship for β in terms of ϕ and OCR
Figure C3 Pile side friction coefficient β in terms of effective friction angle and overconsolidation ratio
C113 Pile Unit End Bearing
C1131 Theoretical Considerations
The evaluation of the end bearing resistance of pile foundations is commonly determined using
limit plasticity theory where
prime Drained loading = C6 119902119887 119873119902 middot 120590119907119900
C-11
Undrained loading 119902119887 = 119873119888 middot 119904119906 C7
where the value of σvo is calculated at the depth z = L and the value of su is the average undrained shear
strength from z = L to a depth z = L + d beneath the pile tip For a circular pile the limit plasticity solution
for undrained loading gives (Vesic 1977) a value Nc = 933 while a deep strip foundation would employ a
value of Nc = 824 For drained loading the expression for Nq for a deep foundation is given by (Vesic
1977)
)]arctan()sin1(tan21[)](tan1[sin1
sin1)tanexp( 2 BLABNq
C8
where A and B are the pile plan dimensions (for a circular pile A = B) and L = pile length For a
circular pile this can be approximated by
asymp 077(120601prime75deg) 119873119902 C9
over a range of effective friction angles 20deg le ϕ le 45deg
C1132 Practical Considerations
For drained loading of sands the full calculated end bearing capacity will never be realized because it
would require the pile to move a distance equal to its diameter This is beyond practical use and
therefore the limit plasticity solutions must be clipped to a fraction of the calculated value Another
reason for using a reduced end-bearing capacity is due to strain incompatibility since the side capacity is
mobilized early while the end-bearing is engaged much later So to have compatible values of Qs and
Qb the qb must be reduced using the following guidelines (Randolph 2003 Mayne 2007)
prime prime C10 119902119887 = 119873119902 ∙ 120590119907119900 ∙ 119891119909
where fx = strain incompatibility factor
fx = 010 for drilled shafts augered cast-in-place and bored piles
fx = 020 for driven low displacement piles (opened-ended steel pipe and H-piles)
fx = 030 for driven high-displacement piles (ie solid piles PSC and closed-ended pipe)
C-12
For undrained loading of circular piles in compression no reduction of end-bearing is necessary
therefore
119902119887 = 933 middot 119904119906 C11
C12 Direct CPT Methods for Pile Capacity
In the direct CPT method the penetrometer readings are scaled directly via specified algorithms to
obtain the pile unit side friction and end-bearing as depicted in Figure C4 As many as 40 different
direct CPT methods have been developed over the past five decades as summarized by Niazi amp Mayne
(2013) Starting circa 1970 many of these early methods relied on hand-recorded data from field
mechanical CPTs where only qc data were obtained at 20 cm intervals or later with mechanical readings
of both qc and fs using special sets of inner and outer rods that recorded vertical load for tip and loads
for tip plus sleeve in alternating increments Also early pile load test data were often obtained from
top-down measurements of load-displacement
C-13
Undrained qb = Nc su
Qside = S (fp DAs)
QTotal = Qs + Qb - Wp
fp = cmck Ko svorsquo tanrsquo
Qbase = qb Ab
Drained qb = Nq svorsquo
Method Two Rational or ldquoIndirectrdquo Method
Method OneldquoDirectrdquo CPT Method
(Scaled Pile)
fp = fctn (soil type pile
type qt fs and u2)
qb = fctn (soil type qt-u2)
qb = unit end bearing
unit sidefriction fp
OCR su Ko t DR rsquo
AXIAL PILE CAPACITY FROM CPT READINGS
Figure C4 Concept of Direct versus Rational Method for CPT evaluation of axial pile capacity
Beginning in the mid-1990s as the modern electric piezocone (CPTu) was implemented the newer
equipment offered improved data and better resolution because of the additional reading of porewater
pressures and correction of raw measured qc to total resistance qt In addition the use of electronic
digital data collection and field computers proved superior in field measurements and recordings As a
consequence several reliable CPT methods for axial pile capacity have been developed Moreover
parallel improvements in full-scale pile load testing occurred and now include modern strain gage
instrumentation digital data recording and automated testing procedures Also the testing can be
single direction (compression or tension) or bi-directional as in the Osterberg cell
C-14
Table C3 provides a selection of recent direct CPT methods that have become available over the
past two decades A number of these (namely ICP NGI UWA and Fugro) were funded by the
offshore industry because of the growth of oil amp gas reserves and windfarm installations thus
necessitating increased concerns on risk probability and reliability in the site investigations for
offshore platforms and design of large driven monopile foundations These direct CPT methods
are often statistically based on large datasets compiled from full-scale load tests made
worldwide (eg Schneider et al 2008)
Table C3 Selection of Direct CPT Methods for Axial Pile Capacity
Method Pile Types Soil Types References CPT
data
Additional
parameters
needed
Unicone Method all types Sands silts
clays
Eslami and
Fellenius
(1997)
Fellenius
(2009)
qt fs u2
KTRI = Kajima
Technical
Research
Institute
all types Sands
mixed clays
Takesue et al
(1998)
fs and u2 No guidance given
on end bearing
resistance
Imperial College
Procedure (ICP)
OE and CE Sands Chow et al
(1997 PhD)
Jardine et al
(2005)
qt Interface friction
angle (δ) from
ring shear tests
Correlated to
mean grain size
(D50)
C-15
OE and CE Clays Jardine et al
(2005)
qt Interface friction
angle (δ)
correlated to PI
Table C3 Continued
NGI Method
(Norwegian
Geotechnical
Institute)
OE and CE Sands Clausen et al
(2005)
qt DR from qt
Clays a Alpha
method
(Karlsrud et al
2005 2012)
b Beta
method
(Karlsrud
2012)
qt for su
qt for
OCR
PI = plasticity
index ()
PI = plasticity
index ()
Fugro Method OE and CE Sands Kolk et al
(2005)
qt
Clays Van Dijk and
Kolk (2011)
qt
OE and CE Sands Lehane et al
(2005)
qt IFR = infilling ratio
related to amount
of plugging
C-16
Purdue LFRD Drilled
shafts
Sands Basu amp
Salgado
(2012)
qt LRFD = load
resistance
factored design
Enhanced Various Sands silts Niazi and qt u2 Based on 330 load
Unicone Method clays and Mayne (2015 and fs tests including
mixed soils 2016) bored augered
jacked and driven
piles
UWA (Univ of
Western
Ra = pile
roughness
Australia) Clays Lehane et al
(2012)
qt Interface friction
angle (δ) from
ring shear tests
and also method
without δ
HKU Method
(Hong Kong
Univ)
OE and CE
(end
resistance
only)
Sands Yu and Yang
(2012)
qt End bearing only
PLR = plug length
ratio
Note PLR can be
estimated from
OE inner diam
Table C3 Continued
Note previously called Marine Technical Directorate (MTD)
C-17
Many of the offshore CPT methods relate to driven piles either in sand or clay as that is
common deep foundation for those purposes Of particular interest are the Unicone and
Modified Unicone Methods since they use all three readings of the piezocone (qt fs and u2)
and address a variety of pile foundation types
C13 Modified UniCone Method
The modified UniCone Method was developed based on a total 330 pile load tests which were
associated with SCPTu data during their site investigations (Niazi and Mayne 2015 2016) This
represents a threefold increase over the original UniCone database that was built upon data
from 106 pile load tests (Eslami amp Fellenius 1997)
For the original UniCone algorithms use is made of the effective cone resistance (qE)
119902119864 = 119902119905 minus 1199062 C12
and a chart of qE vs fs provided an approximate soil classification in five distinct groups as shown by
Figure C5 Later in the modified approach a better delineation of the larger dataset gave soil
subclassifications as indicated by Figure C6
C-18
Figure C5 Soil behavior type using CPT via original UniCone
Figure C6 Soil behavior type using CPT via Modified UniCone
C-19
In the modified approach the 9-zone normalized soil behavioral type (SBTn) is ascertained using CPT
data in conjunction with the Robertson (2009) charts as shown in Figure C7 As discussed in Appendix
A and B the SBTn method uses the normalized cone resistance (Qtn) normalized sleeve friction (Fr())
and CPT material index Ic This permits a much wider range in the calculated pile side friction because fp
is related as a continuous curve with Ic rather than only 5 values that are assigned in the original
scheme
The pile unit side friction (fp) is obtained from qE and the CPT material index Ic using the following
expression at each elevation along the sides of the pile
10(0732 middot 119868119888 minus 3605) fp middot middot C13 = 119902119864 middot 120579119875119879 middot 120579119879119862 120579119877119860119879119864
C-20
Figure C7 Nine-zone soil behavioral soil type using normalized piezocone parameters
(after Robertson 2009 Mayne 2014)
where θPT = coefficient for pile type (θPT = 084 for bored piles 102 for jacked piles 113 for driven
piles) θTC = coefficient for loading direction (θTC = 111 for compression and 085 for tension) and θRATE
= rate coefficient applied to soils in SBT zones 1 through 7 (θRATE = 109 for constant rate of penetration
tests and 097 for maintained load tests) Since CPT provides data at regular intervals of 2 cm to 5 cm
along the sides of the pile the average fp from z = 0 to z = L can be used directly in Equation 1 to obtain
the shaft capacity Qs
C-21
The pile end bearing resistance is obtained from
middot 10(0325middot 119868119888minus1218) 119902119887 = 119902119864 C14
where qE is averaged in the vicinity of the pile tip Figure C8 shows the Modified Unicone Method in
graphical format For sensitive clays of zone 1 please see additional discussions by Niazi amp Mayne
(2016)
Figure C8 Summary of Modified UniCone Method for direct CPT assessment of axial pile
capacity
C-22
C14 Axial Pile Displacement
The movement of pile foundations can be assessed using elastic continuum theory (Poulos amp
Davis 1980 Randolph 2003 Mayne amp Niazi 2017) These relationships have been developed
using finite element analyses boundary elements and analytical closed-form solutions For the
latter approach the top displacement of a rigid pile subjected to a vertical force is shown in
Figure C9 This gives the movement at the top of a pile subjected to either compression or
tension (uplift) loading Also the percentage of axial load transferred to the pile toe is
determined For piles extending through various soil layers the elastic solution can be
implemented by using a set of stacked pile segments each with its own stiffness as
represented by a soil Youngs modulus
For pile groups the use of computer software is recommended Several available programs can handle
pile groups under axial and lateral moment loading such as DEFPIG (Univ Sydney) and PIGLET (Univ
Western Australia) A full listing of pile foundation software is given at the Geotechnical amp
GeoEnvironmental Service Directory wwwggsdcom
Ps
= Sh
aft
Load
sL
t
tEd
IPw
Load Transfer to Base
)]1)((5ln[
)(
)1(1
1
2 vdL
dLFI
E
ECT
211
IF
P
P CT
t
b
Base Load = Pb
Randolph Elastic Solution
Ground Surface
Top Load Pt = Ps + Pb
Rigid Pile Response
wt = displacementd = diameter
L = length
Es = Elastic Soil Modulus
Es0 (at z = 0) surface
EsM (at z = L2)mid-length
EsL (at z = L)full length
Load Transfer to Shaft
E = EsMEsL = Gibson parameterE = 1 for homogeneous case (constant E)E = 05 for pure Gibson case (Es0 = 0)
where FCT = load direction (= 1 compression 0 tension)
21
IF
P
P CT
t
b
z = depth
= Poissons ratio
Tension
Compression
C-23
Figure C9 Elastic continuum solution for axial pile response under compression and tension
loading
C-24
C15 Acknowledgements
The author appreciates the help of Fernando Illingworth of PonsEcuador Meena Viswanth of
GeoSyntecAtlanta and Dr Marco Uzielli of GeoRiskItaly in helping to collect and process the data
Funding support was provided by David Woeller of ConeTec Richmond BC
C-25
- Structure Bookmarks
-
- CONE PENETRATION TEST DESIGN GUIDE FOR STATE GEOTECHNICAL ENGINEERS
- TABLE OF CONTENTS
- LIST OF TABLES
- CHAPTER 1 INTRODUCTION
- CHAPTER 2 DIRECT CPT METHOD FOR SOIL CHARACTERIZATION
- CHAPTER 3 DIRECT CPT METHOD FOR SHALLOW FOUNDATIONS
- CHAPTER 4 DIRECT CPT METHOD FOR DEEP FOUNDATIONS
- REFERENCES
- APPENDIX A
- APPENDIX B
- APPENDIX C
-
Figure 25 Foundation soil formation parameter hs versus CPT material index Ic (Mayne 2017) 60
Figure 26 Diagram of footing profiles for Example 362
Figure 27 CPT data from Northern Minnesota 63
Figure 28 Schematic of foundation design associated with CPT data 64
Figure 29 Soil type for example problem 367
Figure 30 Direct CPT evaluation of axial pile capacity 70
Figure 31 UniCone Method soil behavior type using CPT (Mayne 2017) 71
Figure 32 Modified UniCone Method soil behavior type using CPT (Mayne 2017) 71
Figure 33 Deep foundation end bearing pile diagram74
Figure 34 CPT data from Minnesota for example problem 5 75
Figure 35 Soil layers using ldquorules of thumbrdquo for pile capacity example using direct CPT method 76
Figure 36 Soil layer 1 using SBT method 81
Figure 37 Soil layer 2 using SBT method 82
Figure 38 Soil layer 3 using SBT method 83
Figure 39 Soil layering compared to pilings 85
LIST OF TABLES
Table 1 Geoparameters calculated directly from CPT 4
Table 2 Minor Geoparameters 5
Table 3 Soil type compared to exponent m 10
Table 4 Geoparameters evaluated for Case Example 117
Table 5 Geoparameters 35
Table 6 Bearing capacity defined by soil type61
Table 7 SCPT Results 62
CHAPTER 1 INTRODUCTION
The purpose of this manual is to provide an analysis of soil characterization shallow footings and deep
foundations using direct cone penetration testing (CPT) methods Geotechnical site characterization is
important for evaluating soil parameters that will be used in the analysis and design of foundations
retaining walls embankments and situations involving slope stability Common practice is to determine
these soil parameters through conventional lab and in-situ testing An alternative method uses CPT
readings of cone tip resistance (qt) sleeve friction ( fs) and pore pressure (u2) directly to determine these
parameters such as unit weight effective friction angle undrained shear strength and many others
(Figure 1) Direct CPT methods are provided for these parameters which will be used in the designs for
shallow and deep foundations
Figure 1 Geoparameters determined from CPT
Designing shallow foundations is typically done in a two-part process determining the bearing capacity
and expected settlement (commonly referred to as displacement) of the soil to approximate the
required size and shape of a foundation The older traditional methods are no longer required with
many approaches existing for using CPT directly in the design of shallow foundations Results from these
methods can provide a direct assessment of bearing capacity andor settlement A specific approach to
the direct method has been recommended and tested using a database of 166 full-scale field load tests
(Figure 2) A one-part process is used to scale the measured cone penetrometer readings (ie
measured cone tip resistance sleeve friction and pore water pressure) to obtain the bearing capacity
and settlement of the soil A step-by-step procedure has been created to transition from the CPT data to
bearing capacity with settlement accounted for The steps consist of estimating a design footing width
1
and length while using the process in the Soil Characterization section to determine soil parameters
directly from the CPT data
Figure 2 Conventional method for shallow foundation design compared to direct CPT method
There are upwards of 40 different direct CPT methods that have been developed over the past five
decades to determine the axial compression capacity of a piling foundation Earlier direct CPT methods
relied on hand-recorded information where mechanical-type CPT cone tip resistance data would be
collected at 20 cm intervals
Herein the method recommended for deep foundation design is the Modified UniCone method which
uses all three readings of the modern electronic piezocone penetrometer (CPTu) while addressing a
variety of pile foundation types (Figure 3) The modified UniCone Method is based on a total of 330 pile
load tests (three times the original UniCone database) that were associated with SCPTu data Two
computer software programs have been recommended for analyzing pile movements andor
settlements
2
Figure 3 Direct CPT evaluation of axial pile capacity
3
Symbol Parameter
γt Soil total unit weight
Ic CPT material index
SBT Soil behavior type (SBT)
ʹ σp Preconsolidation stress
YSR Yield stress ratio
ϕʹ Effective friction angle
Ko Lateral stress coefficient
su Undrained shear strength
Dʹ Constrained modulus
Eʹ Drained Youngrsquos modulus
CHAPTER 2 DIRECT CPT METHOD FOR SOIL
CHARACTERIZATION
21 INTRODUCTION
A direct CPT method for determining the value of each of various geoparameters shown in Table 1 is
provided The parameter Ic is used in the derivation of several sequential parameters This section of
the guide may be referred to as these parameters are used in calculations throughout the shallow
foundations and deep foundations sections
Table 1 Geoparameters calculated directly from CPT
4
Kʹ Bulk modulus
ks Subgrade reaction modulus
MR Resilient modulus
Gmax Small-strain shear modulus
k Coefficient of permeability
cv Coefficient of consolidation
Symbol Parameter Equation
ρt Mass Density ρt = γtga
where ga = 98 ms2
σvo Total Stress σvo asymp Σ (γti middot Δzi)
c Effective cohesion ʹ Empirical c asymp 003σp
In clays c asymp 01cu
Other minor parameters can be used in calculations of the geoparameters in Table 1 These minor parameters are provided in Table 2
Table 2 Minor Geoparameters
5
22 SOIL UNIT WEIGHT
The total soil unit weight can be estimated from CPT sleeve friction resistance as shown in Figure 4
(Mayne 2014) This method is not applicable to organic clays diatomaceous soils peats or sensitive
soils
Figure 4 Soil unit weight from CPT sleeve friction
Use Equation 1 to calculate the soils total unit weight
1 120574119905 = 120574119908 ∙ [122 + 015 ∙ ln (100 ∙ 119891119904 + 001)]
120590119886119905119898
Since soil unit weight is required for determining most geoparameters (including Soil Behavior Type)
estimating the soil type of the layers using the ldquorules of thumbrdquo method is a good first step before
determining more precise layering (Figure 5) Once a unit weight is determined for each noticeable layer
change these results can be used in later calculations such as ldquoCPT Material Indexrdquo To use the ldquorules
of thumbrdquo method some helpful guidelines are to assume sands are identified when qt gt 725 psi and u2
asymp uo while the presence of intact clays are prevalent when qt lt 725 psi and u2 gt uo The magnitude of
porewater pressures help to indicate intact clays such as soft (u2 asymp 2middotuo) firm (u2 asymp 4middotuo) stiff (u2 asymp 8middotuo)
and hard (u2 asymp 20middotuo) Fissured overconsolidated clays tend to have negative u2 values such that u2 lt 0
6
Figure 5 Approximate ldquorules of thumbrdquo method using CPT sounding from Wakota Bridge MN
23 CPT MATERIAL INDEX
The development of the CPT material index Ic has improved the initial classification of soil types and
calculation of soil parameters as shown in Table 1 To calculate Ic follow steps 1 and 2
231 Step 1 Normalized Sleeve Friction
Calculate Fr using Equation 2 if not provided with the CPT data gathered
2
7
119891119904 119865119903() = 100 ∙
(119902119905minus120590119907119900)
232 Step 2 Iteration
Iterate using Equations 3-5 to determine Ic by initially using n = 1 to calculate a starting value of Ic The
exponent n is soil-type dependent n = 1 (clays) n asymp 075 (silts) and n asymp 05 (sands) Iteration converges
quickly which is generally after the 3rd cycle
3 (119954119957minus120648119959119952)120648119938119957119950 = 119951 119928119957119951 prime (120648119959119952120648119938119957119950)
4
prime 120590119907119900 119899 = 0381 ∙ 119868119888 + 005 ( ) minus 015 120590119886119905119898
119899 le 10
5 119920119940 = radic[120785 120786120789 minus 119949119952119944119928119957119951]120784 + [120783 120784120784 + 119949119952119944119917119955]120784
24 SOIL BEHAVIOR TYPE (SBT)
Typically soil samples are not taken when using CPT The soil types are then inferred from the qt fs and
u2 readings To determine the types of soil from CPT data follow steps 1 and 2
241 Step 1
To determine the soil layers from CPT results calculate Ic by following the steps under section ldquoCPT
Material Indexrdquo After Ic has been determined through all specified depths use Figure 6 to classify the
type of soil by comparing each Ic value to normalized CPT readings (Fr and Qtn) from Equations 2 and 3
8
Figure 6 SBT zones using CPT Ic
242 Step 2
To determine if any of the soil layers contain ldquosensitive clays and siltsrdquo from zone 1 or ldquovery stiff
overconsolidated (OC) soilrdquo from zones 8 and 9 use Equations 6 and 7 If any soil layers are found within
zone 1 by Equation 6 then caution should be taken as these clays are prone to instability collapses and
difficulties in construction performance Very stiff OC sands to clayey sands of zone 8 (15 lt Fr lt 45)
and very stiff OC clays to silts of zone 9 (Fr gt 45) can be identified by Equation 7
6 Equation 6 errata This is an exponential expression (see Figure 5)
119876119905119899 lt 12119890119909119901(minus14 ∙ 119865119903)
7 119876119905119899 gt 0005(119865119903minus1)minus00003(119865119903minus1)2minus0002
1
Errata See terms and coefficients in Figure 5 above (numbers cannot be rounded off)
After each CPT reading has been assigned a zone from Figure 14 a visual representation can be made to
show the predominant layers by soil types (Figure A5)
25 EFFECTIVE STRESS FRICTION ANGLE
The effective friction angle (ϕ) is used to govern the strength for sands and clays where Equation 8 is
used for sands and Equation 9 is used for clays The value of ϕ for sands is derived from Ic so before ϕ
can be calculated refer to the iteration of Qtn under the ldquoCPT Material Indexrdquo section Once the values
9
Soil Type m
Fissured clays 11
Intact clays 10
Sensitive clays 09
Silt mixtures 085
Silty sands 080
Clean sands 072
Note m may be higher than 11 in fissured clays
of Ic and Qtn are calculated the type of soil can be determined from the section ldquoSoil Behavior Type
SBTrdquo The type of soil will dictate which equation to use for ϕ
Sands
8 120601prime(deg) = 176deg + 110deg log(119876119905119899)
Clays
9 119902119905minus120590119907119900 120601prime(deg) = 295deg ∙ 119861119902
0121 ∙ [0256 + 0336 ∙ 119861119902 + log ( )] prime 120590119907119900
Where = 119861119902 (1199062 minus 119906119900)(119902119905 minus 120590119907119900)
26 STRESS HISTORY
Determining the stress history can be characterized by an apparent yield stress ratio of the form
10 prime 120590119901
119884119878119877 = prime 120590119907119900
Where σp is defined as the preconsolidation stress or effective yield stress (Equation 10) The YSR is the
same equation as the more common overconsolidation ratio (OCR) but is now generalized to
accommodate mechanisms of preconsolidation such as ageing desiccation repeated cycles of wetting-
drying repeated freeze thaw cycles and other factors
11 prime prime 120590119901 = 120590119910 = 033(119902119905 minus 120590119907119900)119898prime
Where σp σy σvo and qt have units of kPa and the value of m depends on soil type with typical values shown in Table 3
Table 3 Soil type compared to exponent m
10
The value of m for non-fissured soils and inorganic clays and silts is derived from Ic (Figure 7) so before
m can be calculated (Equation 12) refer to the iteration of Ic under the ldquoCPT Material Indexrdquo section
Determine Ic for all soil layers before calculating m
12 028
119898prime = 1 minus 25 119868119888 1+( frasl ) 265
Once m is known for all soil layers the YSR can be determined
Figure 7 Yield stress exponent compared to CPT material index
A limiting value of YSR can be reached for clays and sands It can be calculated using Equation 13
13 (1 sin(emptyprime))
(1+sin(emptyprime)) 119884119878119877119897119894119898119894119905 = [ ]
(1minussin(emptyprime))2
27 LATERAL STRESS COEFFICIENT
The lateral stress coefficient Ko = σhoσvo commonly referred to as the at-rest condition is used to
represent the horizontal geostatic state of soil stress Ko can be calculated using Equation 14 but
11
14
119870119900 = (1 minus sin(emptyprime)) ∙ 119884119878119877sin( emptyprime)
sections such as ldquoStress Historyrdquo and ldquoEffective Stress Friction Anglerdquo will need to be referred to for
determining parameters ϕ and YSR
A maximum value for Ko can be determined by Equation 15
15 (1+sin(emptyprime))
119870119900119898119886119909 = = 1199051198861198992(45deg + emptyprime2) (1minussin(emptyprime))
28 UNDRAINED SHEAR STRENGTH
Loading on soils can result in fully drained partially drained or fully undrained conditions Sands
typically produce drained cases due to their high permeability but exceptions may occur in loose sands
during fast loading where the water does not have sufficient time to dissipate Clays exhibit low
permeability and thus often result in undrained loading cases when a load is applied quickly For soft-
firm clays the undrained shear strength (su) can be determined from CPT via Equation 16 where the
value of the bearing factor Nkt can be taken as 12
16 119902119905minus120590119907119900 119904119906 =
119873119896119905
In the case of remolded undrained shear strength from CPT su asymp fs
29 GROUND STIFFNESS AND SOIL MODULI
Determining the grounds stiffness can be measured from geoparameters such as the constrained
modulus (D) drained Youngrsquos modulus (E) bulk modulus (K) subgrade reaction modulus (ks) resilient
modulus (MR) and small-strain shear modulus (Gmax)
17 119863 prime asymp 5 ∙ (119902119905 minus 120590119907119900)
18
12
119864 prime = 119863 prime
11
19 119864prime
119870prime = [3∙(1minus2119907prime)]
The subgrade modulus (ks) is a combination of soil-structural properties which creates a parameter that
depends on the ground stiffness and the size of the loaded element
20 119864prime
119896119904 = [119889∙(1minus1199072)]
The resilient modulus MR applies to pavement analysis and design and can be calculated using Equation
21 where qt and fs are in MPa
21 053 119872119877 = (146119902119905 + 135511989111990414 + 236)244
The small strain shear modulus Gmax is a representation of the initial stiffness of all soils and rocks
Graphically it is the beginning portion of all stress-strain-strength curves for geomaterials Use Equation
22 to determine Gmax
22 119866119898119886119909 = 120588119905 ∙ 1198811199042
Where Vs = shear wave velocity (Equation 23) as measured by seismic cone penetration tests (SCPT) If
only cone penetration tests (CPT) or piezocone (CPTu) data are available the shear wave velocity may
be estimated from
23 03 119891119904 119881119904 (119898frasl119904) = [101 ∙ log(119902119905) minus 114]167 ∙ (100 ∙ )
119902119905
Where qt and fs have units of (kPa)
210 COEFFICIENT OF CONSOLIDATION
The coefficient of consolidation (cv) controls the rate that foundation and embankment settlements
occur By using results of CPT dissipation tests that measure the change in u2 readings over time the
value of cv can be determined Using Equation 24 cv can be determined based on piezocone dissipation
13
curves The equation below requires an estimate of the in-situ rigidity index (IR) of the soil If results of
SCPTU are available then IR may be determined from Gmax and qt per equation A38
24 0030∙(119886119888)2∙(119868119877)075
119888119907 = 11990550
Where ac = penetrometer radius (178 cm for 10-cm2 cone 220 cm for 15-cm2 cone) t50 = time to reach 50 dissipation IR = Gsu = undrained rigidity index G can be determined from the ldquoGround Stiffness and Soil Modulirdquo section
211 HYDRAULIC CONDUCTIVITY
The hydraulic conductivity also known as the coefficient of permeability (k) expresses the flow
characteristics of soils and has units of cms or feetday One method of calculation would be to use
Equation 25 where cv would need to be determined from ldquoCoefficient of Consolidationrdquo and D would
need to be determined from ldquoGround Stiffness and Soil Modulirdquo
25 119888119907∙120574119908 119896 =
119863prime
An alternative approach developed for soft normally-consolidated soils (Figure 8) is shown in Equation
26 where t50 (sec) values are used directly in assessing k in (cms)
26
14
125
119896 asymp ( 1
) 251∙11990550
Figure 8 k vs dissipation time for 50 consolidation (Mayne 2017)
212 EXAMPLE PROBLEMS
2121 Example 1 Direct CPT Methods for Geoparameters on Sands
Several geoparameters need to be determined based on the given CPT data collected for the South
Abutment of a bridge in Benton County MN (Figure 9) The groundwater table (GWT) was measured at
17 feet Determine all the geoparameters found in Table 4 at depths of 0 feet to 30 feet All sand layers
can be assumed ldquodrainedrdquo with ν = 02
15
Figure 9 CPT data from Benton County Minnesota for example problem 1
16
Table 4 Geoparameters evaluated for Case Example 1
Symbol Parameter Symbol Parameter
γt Soil total unit weight Ko Lateral stress coefficient
Ic CPT material index Dʹ Constrained modulus
SBT Soil behavior type (SBT) Eʹ Drained Youngrsquos modulus ʹ σp Preconsolidation stress Kʹ Bulk modulus
YSR Yield stress ratio MR Resilient modulus
ϕʹ Effective friction angle Gmax Small-strain shear modulus
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 10 Soil layers using ldquorules of thumbrdquo
γw = 6224 pcf
17
Layer 1
From Figure 10 fs = 17 psi taken as a representative value of the layer
17 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1204 pcf
145 psi
Layer 2
From Figure 10 fs = 17 psi taken as a representative value of the layer
17psi
γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1204 pcf 145 psi
Layer 3
From Figure 10 fs = 12 psi taken as a representative value of the layer
12 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 4
From Figure 10 fs = 7 psi taken as a representative value of the layer
CPT Material Index
7 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1121 pcf
145 psi
Layer 1
From Figure 10 fs = 17 psi and qc = 3500 psi
18
qt = qc + u2 ∙ (1 minus a) = 3500 psi + 3 psi ∙ (1 minus 08) = 3501 psi
= γt ∙ 6 feet = 1204 pcf ∙ 6 feet = 7225 psf = 50 psi σvo
fs 17 psi Fr() = 100 ∙ = 100 ∙ = 049
(qt minus σvo) (3501 psi minus 50 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 5 psi minus 0 psi = 50 psi
(qt minus σvo)σatm (3501 psi minus 50 psi )145 psi = = Qtn prime )n (σvoσatm (50 psi145 psi)n
prime σvo 50 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(049)]2
Qtn = 35329 n = 036 lt 10 Ic = 13
Layer 2
From Figure 10 fs = 17 psi and qc = 3500 psi
19
qt = qc + u2 ∙ (1 minus a) = 3500 psi + 0 psi ∙ (1 minus 08) = 3500 psi
= 7225 psf + γt ∙ 6 feet = 7225 psf + 1204 pcf ∙ 6 feet = 1445 psf = 100 psi σvo
fs 17 psi Fr() = 100 ∙ = 100 ∙ = 049
(qt minus σvo) (3500 psi minus 100 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 100 psi minus 0 psi = 100 psi
(qt minus σvo)σatm (3500 psi minus 100 psi )145 psi = = Qtn prime )n (σvoσatm (100 psi145 psi)n
prime σvo 100 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
47 minus log(Qtn)]2 + [122 + log(049)]2
Ic = radic[3
Qtn = 27947 n = 041 lt 10 Ic = 14
Layer 3
From Figure 10 fs = 12 psi and qc = 1500 psi
20
qt = qc + u2 ∙ (1 minus a) = 1500 psi + 0 psi ∙ (1 minus 08) = 1500 psi
= 1445 psf + γt ∙ 11 feet = 1445 psf + 1172 pcf ∙ 11 feet = 2734 psf = 190 psi σvo
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 081
(qt minus σvo) (1500 psi minus 190 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = γwater ∙ (z minus 119911119908) = 6224 pcf ∙ (23 feet minus 17 feet) = 3734 psf = 99 psi
prime σvo = σvo minus uo = 190 psi minus 99 psi = 164 psi
(qt minus σvo)σatm (1500 psi minus 190 psi )145 psi = = Qtn prime )n (σvoσatm (164 psi145 psi)n
prime σvo 164 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(081)]2
Qtn = 947 n = 06 lt 10 Ic = 19
21
Layer 4
From Figure 10 fs = 7 psi and qc = 1200 psi
qt = qc + u2 ∙ (1 minus a) = 1200 psi + 0 psi ∙ (1 minus 08) = 1200 psi
= 2734 psf + γt ∙ 11 feet = 2734 psf + 1121 pcf ∙ 7 feet = 3519 psf = 244 psi σvo
fs 7 psi Fr() = 100 ∙ = 100 ∙ = 060
(qt minus σvo) (1200 psi minus 244 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = γwater ∙ (z minus 119911119908) = 6224 pcf ∙ (30 feet minus 17 feet) = 8091 psf = 130 psi
prime σvo = σvo minus uo = 244 psi minus 130 psi = 188 psi
(qt minus σvo)σatm (1200 psi minus 244 psi )145 psi = = Qtn prime )n (σvoσatm (188 psi145 psi)n
prime σvo 188 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(060)]2
22
Qtn = 686 n = 064 lt 10 Ic = 19
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic Qtn and Fr the first layer is defined as a ldquoDrained Gravelly Sandrdquo from
Figure 11
Figure 11 Soil layer 1 using SBT method
23
Layer 2
Based on values of Ic Qtn and Fr the second layer is defined as a ldquoDrained Sandrdquo from Figure
12
Figure 12 Soil layer 2 using SBT method
24
Layer 3
Based on values of Ic Qtn and Fr the third layer is defined as a ldquoDrained Sandrdquo from Figure 13
Figure 13 Soil layer 3 using SBT method
25
Layer 4
Based on values of Ic Qtn and Fr the fourth layer is defined as a ldquoDrained Sandrdquo from Figure 14
Figure 14 Soil layer 4 using SBT method
Effective Stress Friction Angle
Layer 1
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(3533) = 456deg
Layer 2
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(2795) = 445deg
26
Layer 3
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(947) = 393deg
Layer 4
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(686) = 378deg
Stress History
Layer 1
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (13frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(24139 kPa minus 345 kPa)072 = 4717 kPa
= 684 psi
prime σp 684 psi YSR = = = 136
primeσvo 50 psi
Layer 2
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + (14 1 + ( frasl ) frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(34132 kPa minus 689 kPa)072 = 605 kPa
= 877 psi
prime σp 877 psi YSR = = = 87
primeσvo 100 psi
27
Layer 3
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + (19 1 + ( frasl ) frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(10342 kPa minus 131 kPa)072 = 254 kPa
= 368 psi
prime σp 368 psi YSR = = = 22
primeσvo 164 psi
Layer 4
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (19frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(8274 kPa minus 168 kPa)072 = 215 kPa = 312 psi
prime σp 312 psi YSR = = = 17
primeσvo 188 psi
Lateral Stress Coefficient
Layer 1
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(456deg)) ∙ 136sin( 456deg) = 18
Layer 2
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(445deg)) ∙ 87sin( 445deg) = 14
28
Dprime 17480 psi
Eprime = = = 15890 psi 11 11
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(393deg)) ∙ 22sin( 393deg) = 06
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(378deg)) ∙ 17sin( 378deg) = 05
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3501 psi minus 50 psi) = 17480 psi
Eprime 15890 psi
Kprime = = = 8828 psi [3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Layer 3
Layer 4
Ground Stiffness and Soil Moduli
Layer 1
Values of qt and fs need to be in MPa
MR = (146(241 MPa)053 + 1355(012 MPa)14 + 236)244 = 3416 MPa = 49575 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1172 kPa m Vs (mfrasls) = [101 ∙ log(24139 kPa) minus 114]167 ∙ (100 ∙ ) = 2745
24139 kPa s
Vs = 9012 fts
29
γt 1204 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 37 ∙ (9012 ) = 304 ∙ 106psf = 21000 psi s
Layer 2
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3500 psi minus 100 psi) = 17480 psi
Dprime 17480 psi Eprime = = = 15890 psi
11 11
Eprime 15890 psi Kprime = = = 8828 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(241 MPa)053 + 1355(012 MPa)14 + 236)244 = 3415 MPa = 49562 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1172 kPa m Vs (mfrasls) = [101 ∙ log(24133 kPa) minus 114]167 ∙ (100 ∙ ) = 2745
24133 kPa s
Vs = 9012 fts
γt 1204 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
30
2 ft Gmax = ρt ∙ Vs
2 = 37 ∙ (9012 ) = 304 ∙ 106psf = 21000psi s
Layer 3
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (1500 psi minus 190 psi) = 7405 psi
Dprime 7405 psi Eprime = = = 6732 psi
11 11
Eprime 6732 psi Kprime = = = 3740 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(103 MPa)053 + 1355(008 MPa)14 + 236)244 = 1506 MPa = 21854 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 827 kPa m Vs (mfrasls) = [101 ∙ log(10343 kPa) minus 114]167 ∙ (100 ∙ ) = 2611
10343 kPa s
Vs = 8566 fts
γt 1172 pcf slug ρt = = = 36
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 36 ∙ (8566 ) = 27 ∙ 106psf = 19000 psi s
Layer 4
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (1200 psi minus 244 psi) = 5878 psi
31
Dprime 5878 psi Eprime = = = 5343 psi
11 11
Eprime 5343 psi Kprime = = = 2969 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(83 MPa)053 + 1355(005 MPa)14 + 236)244 = 165 MPa = 16901 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 483 kPa m Vs (mfrasls) = [101 ∙ log(8274 kPa) minus 114]167 ∙ (100 ∙ ) = 2243
8274 kPa s
Vs = 7360 fts
γt 1121 pcf slug ρt = = = 35
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 35 ∙ (7360 ) = 389 ∙ 106psf = 13000 psi s
32
2122 Example 2 Direct CPT Methods for Geoparameters on Clay
Several geoparameters need to be determined based on the given CPT data collected for the South
Abutment (Figure 15) The groundwater table (GWT) was measured at 60 feet Determine all the
geoparameters found in Table 5 at depths of 0 feet to 42 feet All sand layers can be assumed ldquodrainedrdquo
ν = 02 All clay layers can assume ν = 049
33
Figure 15 CPT data from Minnesota for example problem 2
34
Table 5 Geoparameters
Symbol Parameter Symbol Parameter
γt Soil total unit weight Eʹ Drained Youngrsquos modulus
Ic CPT material index Kʹ Bulk modulus
SBT Soil behavior type (SBT) MR Resilient modulus
ʹ σp Preconsolidation stress Gmax Small-strain shear modulus
YSR Yield stress ratio su Undrained shear strength
ϕʹ Effective friction angle cv Coefficient of consolidation
Ko Lateral stress coefficient k Hydraulic conductivity
Dʹ Constrained modulus
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
35
Figure 16 Soil layers using ldquorules of thumbrdquo
Unit weight of water γw = 6224 pcf
Layer 1
From Figure 16 fs = 13 psi taken as a representative value of the layer
13 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1179 pcf
145 psi
Layer 2
From Figure 16 fs = 12 psi taken as a representative value of the layer
12 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 3
From Figure 16 fs = 20 psi taken as a representative value of the layer
36
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
Layer 4
From Figure 16 fs = 2 psi taken as a representative value of the layer
2 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1004 pcf
145 psi
CPT Material Index
Layer 1
From Figure 10 fs = 13 psi and qc = 3000 psi
qt = qc + u2 ∙ (1 minus a) = 3000 psi + 0 psi ∙ (1 minus 08) = 3000 psi
σvo = γt ∙ 2 feet = 1179 pcf ∙ 2 feet = 235 psf = 16 psi
fs 13 psi Fr() = 100 ∙ = 100 ∙ = 043
(qt minus σvo) (3000 psi minus 16 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 16 psi minus 0 psi = 16 psi
(qt minus σvo)σatm (3000 psi minus 16 psi )145 psi = = Qtn prime )n (16 psi145 psi)n (σvoσatm
prime σvo 16 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(043)]2
37
Qtn = 4126 n = 032 lt 10 Ic = 121 therefore sand
Layer 2
From Figure 10 fs = 12 psi and qc = 250 psi
qt = qc + u2 ∙ (1 minus a) = 250 psi + 10 psi ∙ (1 minus 08) = 252 psi
σvo = 235 psf + γt ∙ 30 feet = 235 psf + 1172 pcf ∙ 30 feet = 3751 psf = 260 psi
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 53
(q minus σ ) (252 psi minus 260 psi) t vo
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 260 psi minus 0 psi = 260 psi
(qt minus σvo)σatm (250 psi minus 260 psi )145 psi = = Qtn prime (σvoσatm)n (260 psi145 psi)n
38
prime σvo 260 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(53)]2
Qtn = 87 n = 10 le 10 Ic = 32 therefore clay
Layer 3
From Figure 10 fs = 20 psi and qc = 4000 psi
qt = qc + u2 ∙ (1 minus a) = 4000 psi + 8 psi ∙ (1 minus 08) = 40016 psi
σvo = 3751 psf + γt ∙ 2 feet = 3751 psf + 1219 pcf ∙ 2 feet = 3995 psf = 277 psi
fs 16 psi Fr() = 100 ∙ = 100 ∙ = 054
(qt minus σvo) (30016 psi minus 277 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 277 psi minus 0 psi = 277 psi
(qt minus σvo)σatm (30016 minus 277)145 psi = = Qtn prime (σvoσatm)n (277 psi145 psi)n
39
primeσvo 277 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(054)]2
Qtn = 1962 n = 052 le 10 Ic = 15 therefore sand
Layer 4
From Figure 10 fs = 2 psi and qc = 250 psi
qt = qc + u2 ∙ (1 minus a) = 250 psi + 0 psi ∙ (1 minus 08) = 250 psi
= 3994 psf + γt ∙ 8 feet = 3994 psf + 1004 pcf ∙ 8 feet = 4798 psf = 333 psi σvo
fs 2 psi Fr() = 100 ∙ = 100 ∙ = 092
(qt minus σvo) (250 psi minus 333 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 333 psi minus 0 psi = 333 psi
(qt minus σvo)σatm (250 psi minus 333 psi )145 psi = = Qtn prime (σvoσatm)n (333 psi145 psi)n
40
prime σvo 333 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(092)]2
Qtn = 65 n = 10 lt 10 Ic = 29 therefore clayey silt
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic Qtn and Fr the first layer is defined as a ldquoDrained Sandrdquo from Figure 17
41
Figure 17 Soil layer 1 using SBT method
Layer 2
Based on values of Ic Qtn and Fr the second layer is defined as a ldquoUndrained Clayrdquo from Figure
18
42
Figure 18 Soil layer 2 using SBT method
43
Layer 3
Based on values of Ic Qtn and Fr the third layer is defined as a ldquoDrained Sandrdquo from Figure 19
Figure 19 Soil layer 3 using SBT method
44
Layer 4
Based on values of Ic Qtn and Fr the fourth layer is defined as a ldquoUndrained Silty Mixrdquo from
Figure 20
Figure 20 Soil layer 4 using SBT method
Effective Stress Friction Angle
Layer 1 (sand)
45
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(4126) = 464deg
Layer 2 (clay)
qt minus σvo
ϕprime(deg) = 295deg ∙ Bq0121 ∙ [0256 + 0336 ∙ Bq + log ( )]
primeσvo
Where
(u2 minus uo) (10 psi minus 0 psi) Bq = = = 004
(qt minus σvo) (252 psi minus 260 psi)
252 psi minus 260 psi ϕprime(deg) = 295deg ∙ 00440121 ∙ [0256 + 0336 ∙ 0044 + log ( )]
260 psi
ϕprime(deg) = 245deg
Layer 3 (sand)
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(1962) = 428deg
Layer 4 (clay)
qt minus σvo
ϕprime(deg) = 295deg ∙ Bq0121 ∙ [0256 + 0336 ∙ Bq + log ( )]
primeσvo
Where
(u2 minus uo) (5 psi minus 0 psi) Bq = = = 002
(qt minus σvo) (251 psi minus 333 psi)
252 psi minus 333 psi ϕprime(deg) = 295deg ∙ 0020121 ∙ [0256 + 0336 ∙ 002 + log ( )]
333 psi
ϕprime(deg) = 265deg
Stress History
46
Layer 1
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (12frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(3000 psi minus 33)072 = 1051 psi
prime σp 1051 psi YSR = = = 642
primeσvo 16 psi
Layer 2
028 028
prime m = 1 minus = 1 minus = 099 25 25 1 + (
Icfrasl ) 1 + (32frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(252 psi minus 260)099 = 734 psi
prime σp 734 psi YSR = = = 28
primeσvo 260 psi
Layer 3
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + ( frasl ) 1 + (15frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(4002 psi minus 277)072 = 1288 psi
47
prime σp 1288 psi YSR = = = 46
primeσvo 277 psi
Layer 4
028 028
prime m = 1 minus = 1 minus = 097 25 25 Ic 1 + ( frasl ) 1 + (29frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(250 psi minus 333)097 = 626 psi
prime σp 626 psi YSR = = = 19
primeσvo 333 psi
Lateral Stress Coefficient
Layer 1
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(464deg)) ∙ 642sin( 464deg) = 56
Layer 2
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(245deg)) ∙ 28sin( 245deg) = 090
Layer 3
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(428deg)) ∙ 46sin( 428deg) = 091
Layer 4
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(266deg)) ∙ 19sin( 266deg) = 073
48
Ground Stiffness and Soil Moduli
Layer 1
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3000 psi minus 16 psi) = 14991 psi
Dprime 14991 psi Eprime = = = 13629 psi
11 11
Eprime 13629 psi Kprime = = = 7571 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(207 MPa)053 + 1355(009 MPa)14 + 236)244 = 2816 MPa = 40873 psi
fs
03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 896 kPa m Vs (mfrasls) = [101 ∙ log(20685 kPa) minus 114]167 ∙ (100 ∙ ) = 2564
20685 kPa s
Vs = 8412 fts
γt 1179 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
2 ft 2 Gmax = ρt ∙ Vs = 37 ∙ (8412 ) = 260 ∙ 106psf = 17992 psi
s
Layer 2
49
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (252 psi minus 260 psi) = 11298 psi
Dprime 11298 psi Eprime = = = 10271 psi
11 11
Eprime 10271 psi Kprime = = = 17118 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(049))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(17 MPa)053 + 1355(008 MPa)14 + 236)244 = 443 MPa = 64309 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 827 kPa m Vs (mfrasls) = [101 ∙ log(17375 kPa) minus 114]167 ∙ (100 ∙ ) = 2646
17375 kPa s
Vs = 8680 fts
γt 1172 pcf slug ρt = = = 36
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 36 ∙ (8680 ) = 274 ∙ 106psf = 19036 psi s
Layer 3
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (4002 psi minus 277 psi) = 19869 psi
Dprime 19869 psi Eprime = = = 18063 psi
11 11
50
Eprime 18063 psi Kprime = = = 10035 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(276 MPa)053 + 1355(014 MPa)14 + 236)244 = 4017 MPa = 58309 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1379 kPa m Vs (mfrasls) = [101 ∙ log(27591 kPa) minus 1379]167 ∙ (100 ∙ ) = 2854
27591 kPa s
Vs = 9363 fts
γt 121 pcf slug ρt = = = 38
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 38 ∙ (9363 ) = 33 ∙ 106psf = 23050 psi s
Layer 4
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (250 psi minus 333 psi) = 1088 psi
Dprime 1088 psi Eprime = = = 989 psi
11 11
Eprime 989 psi Kprime = = = 16490 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(049))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
51
MR = (146(17 MPa)053 + 1355(001 MPa)14 + 236)244 = 360 MPa = 52189 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 138 kPa m Vs (mfrasls) = [101 ∙ log(17238 kPa) minus 114]167 ∙ (100 ∙ ) = 1545
17238 kPa s
Vs = 5069 fts
γt 1004 pcf slug ρt = = = 31
ga 322 ftfrasls2 ft3
2 ft 2 Gmax = ρt ∙ Vs = 31 ∙ (5069 ) = 80 ∙ 105psf = 5565 psi
s
Undrained Shear Strength
Layer 2
qt minus σv0 250 psi minus 26 psi su = = = 187 psi
12 12
Layer 4
qt minus σv0 250 psi minus 333 psi su = = = 181 psi
12 12
Coefficient of Consolidation
52
Example dissipation data are shown in Figure 21 Example calculations are provided
Figure 21 Dissipation t50 data
Layer 1
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (7197)075
cv = = = 359 cmsec 1 sec t50
Layer 2
53
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (94392)075
cv = = = 0008 cmsec 3000 sec t50
Layer 3
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (8303)075
cv = = = 665 cmsec 06 sec t50
Layer 4
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (26714)075
cv = = = 087 cmsec 11 sec t50
Hydraulic Conductivity
Layer 1
cm 1 psi 359 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 10 ∙ 10minus4 cmsec Dprime 14984 psi
Layer 2
cm 1 psi 0008 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 73 ∙ 10minus7 cmsec Dprime 11379 psi
Layer 3
cm 1 psi 665 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 20 ∙ 10minus5 cmsec Dprime 148650 psi
54
Layer 4
cm 1 psi 087 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 18 ∙ 10minus4 cmsec Dprime 10861 psi
55
CHAPTER 3 DIRECT CPT METHOD FOR SHALLOW
FOUNDATIONS
31 PROCEDURE
Shallow foundation analysis is typically done in a two-part traditional procedure The traditional
techniques are no longer required as a direct CPT method for square rectangular and circular shallow
footings is available (Figure 22) This process has the soil types grouped into four main categories sands
silts fissured clays and intact clays When determining soil types for each design it is believed footings
on sands and silts act in a fully drained manner while intact clays act in an undrained manner under
conditions of constant volume In order to determine the vertical stress-displacement-capacity of
square rectangular and circular shallow footings follow the steps provided towards the solution given
by Equation 27
Figure 22 Direct CPT method for shallow foundations
Equation 27 may be used to calculate all footing stresses from zero to the bearing capacity (qmax) To
calculate qmax for a sized footing of width (B) length (L) and thickness (t) follow steps 1 through 7
provided The settlement (s) can be determined after the calculation of qmax by simply rearranging
56
Equation 27 where the allowable stress (qallow) is defined as qmax divided by the factor of safety (FS) For
shallow footings a FS value of 3 is commonly used in geotechnical engineering
120782120787 minus120782120785120786120787 119956 119923 119954119950119938119961 = 119945119956 ∙ 119954119957119951119942119957 ∙ (119913
) ∙ (119913
) 27 119950119938119961
2 119902119898119886119909 0345 1 frasl119865119878 119871 119904 = 119861 ∙ [ ∙ ∙ ( ) ] 28
119861 ℎ119904 119902119905119899119890119905
Where hs = the foundation soil formation parameter qtnet = the net corrected cone tip resistance
311 Step 1 Estimating Footing Dimensions
In Minnesota frost heave can have devastating effect on a shallow foundations It is common practice to
place a foundation bearing elevation below the expected maximum frost depth (roughly 45 to 6 feet
below ground level) With this assumption estimate a footing size (B x L) for design to obtain
representative data from CPT roughly 15middotB below the foundation depth (Df) as shown in Figure 23 The
CPT data collected will consist of
cone tip resistance (qc)
measured porewater pressure acting behind the cone tip (u2) as shown in Figure 24
sleeve friction (fs)
57
Figure 23 Direct CPT method introduction
Figure 24 Differentiation of porewater pressure measurement locations (Lunne et al 1997)
312 Step 2 Soil Characterization
The soil behavior type (SBT) and CPT material index (Ic) govern the value of the formation factor hs The
first step is following the steps in Soil Unit Weight to determine soil layering and γt values for each
layer To determine a representative unit weight (γsoil) for all the layers in the range of Df to 15∙B below
58
Df the user will need to use their engineering judgement on the definition of ldquorepresentative unit
weightrdquo This single value of unit weight is used in further calculations such as the total vertical soil stress (σvo) from Equation 29 and effective vertical stress (σvo) from Equation 30 Values of σvo and σvo
will need to be calculated at 15middotB below Df
120648119959119952 = sum(120632119956119952119946119949 ∙ 119963) 29
prime 120648119959119952 = 120648119959119952 minus 119958119952 30
Where z = Df +15middotB uo=γwatermiddot(z-zw)
The qc will need to be corrected using Equation 31 in the case of fine-grained soils that develop excess porewater pressure during cone penetration These values will be used in the following calculations
119954119957 = 119954119940 + 119958120784 ∙ (120783 minus 119938) 31
Where 119860119899 a = cone area ratio = eg MnDOT commonly uses a = 08 119860119888
119860119899 = cross-sectional area of load cell or shaft 119860119888 = projected area of the cone
The cone area ratio is determined based on the type of piezocone tip used during in-situ field testing
Manufacturer specifications should provide the measured net area ratio (a) for the particular cone
penetrometer as determined by calibration in a pressurized triaxial chamber
313 Step 2a Foundation soil formation parameter
With the representative value γsoil continue to follow the steps in CPT Material Index through Soil
Behavior Type (SBT) to better define the type of soil at depth 15∙B below Df Use the value of Ic
calculated at depth 15∙B below Df to calculate hs with Equation 36 This parameter is based on the soil
type with typical values shown in Figure 25 The data on silts and sands are considered fully drained
whereas the fissured clay subset may be partially drained to undrained
59
23 ℎ119904 = 28 minus
119868119888 15 32
1+( ) 24
Figure 25 Foundation soil formation parameter hs versus CPT material index Ic (Mayne 2017)
314 Step 3 Soil elastic modulus and Poissonrsquos ratio
Details about the soil such as its elastic modulus (Es) and Poissonrsquos ratio (ν) will be needed for further
calculations A representative value for the Es can be determined from in-situ field tests Values of ν can
be assigned as 02 for drained sands and as 05 for undrained loading cases involving clays (Jardine et al
1985 Burland 1989)
315 Step 4 Net cone tip resistance
Calculate qtnet the mean value of net cone tip resistance 15middotB below the foundation bearing elevation
using Equation 33
33 119902119905119899119890119905 = 119902119905 minus 120590119907119900
60
316 Step 5 Bearing capacity of the soil
Use the assumed B and L calculated hs and qtnet to determine the soils bearing capacity (qmax) from
Equation 34 Use Table 6 to calculate qmax by using the maximum allowable settlement ratio (sB)max
correlating to the soil type If hs is in between the given values interpolate to acquire (sB)max
Table 6 Bearing capacity defined by soil type
Type of Soil hs (sB)max
Clean Sands 058 12
Silts 112 10
Fissured Clays 147 7
Intact Clays 270 4
05 minus0345 119904 119871 119902119898119886119909 = ℎ119904 ∙ 119902119905119899119890119905 ∙ (119861
) ∙ (119861
) 34 119898119886119909
317 Step 6 Settlement
Settlement can be calculated directly using the results from Equation 35 A FS equal to 3 is common in
foundation engineering
2 119902119898119886119909 0345 1 frasl119865119878 119871 119904 = 119861 ∙ [ ∙ ∙ ( ) ] 35
119861 ℎ119904 119902119905119899119890119905
318 Step 7 Final Check
Check that the applied stress (q) is less than qmax using Equation 36 Repeat the process again with a
new B andor new L if q gt qmax
05 119904 ) 119902 = ℎ119904 ∙ 119902119905119899119890119905 ∙ (119861
minus0345 119871 ∙ ( )
11986136
61
32 EXAMPLE PROBLEMS
321 Example 3 Direct CPT Method on Sands
A footing size needs to be determined based on the given CPT data collected for the South Abutment
(Figure 27) The footing stress (q) was determined to be 8000 psf Estimate a footing size (B x L) and
determine the bearing capacity of the foundation (qmax) using the direct CPT method provided Also
determine the expected settlement based on the calculated bearing capacity
Figure 26 Diagram of footing profiles for Example 3
The soil elastic modulus determined from seismic CPT (SCPT) are shown in Table 7
Table 7 SCPT Results
Depth (feet) Es (tsf)
Bottom of layer
3 557
6 433
9 557
12 695
17 590
22 501
27 571
32 505
62
Figure 27 CPT data from Northern Minnesota
Solution
Step 1 Assume the footing is placed below the frost depth of 6 feet
Estimate L L = 50 feet = 600 inches
63
Estimate B B = 12 feet = 144 inches
Estimate footing thickness t t = 2 feet
Df = 6 feet = 72 inches
Df + 15middotB = 24 feet = 288 inches
Step 2 Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 28 Schematic of foundation design associated with CPT data
Layer 1
From Figure 28 fs = 20 psi taken as a representative value of the layer
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
64
Layer 2
From Figure 28 fs = 20 psi taken as a representative value of the layer
20psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
Layer 3
From Figure 28 fs = 8 psi taken as a representative value of the layer
8 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1134 pcf
145 psi
Using the unit weights calculated between Df and 15B determine a representative unit weight of the soil to calculate the total and effective soil stresses Based on layer 3 being the weakest supporting layer γsoil = 113 pcf
σvo = γsoil ∙ (119863119891 + 15 ∙ B) = 113 lbsfraslft3 ∙ 24 feet = 2721 psf = 189 psi
prime σvo = σvo minus uo = 2721 psf minus 0 psf = 2721 psf = 189 psi
Calculate the cone tip resistance between Df and Df + 15B below the foundation depth
qt = qc + u2 ∙ (1 minus a) = 1250 psi + 0 psi ∙ (1 minus 08) = 1000 psi
Step 2a CPT Material Index
Using Figure 28 the sleeve friction at 15B below the foundation depth is about 16 psi
65
fs 16 psi Fr() = 100 ∙ = 100 ∙ = 13
(qt minus σvo) (1250 psi minus 189 psi)
Iterate to solve for Ic Steps are not shown for brevity
(qt minus σvo)σatm (1250 psi minus 189 psi )145 psi = = Qtn prime (σvoσatm)n (189 psi145 psi)n
prime σvo 189 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Qtn = 703 n = 072 lt 10 Ic = 210
66
Figure 29 Soil type for example problem 3
Use the Ic value to determine the foundation soil formation parameter
Calculate the foundation soil formation parameter The tip resistance is about 1250 psi at 24 feet and
the cone area ratio was determined to be 08 from information provided by the manufacturer
23 23 hs = 28 minus = 28 minus = 078 = SandSilt 15 15 Ic 210
1 + ( ) 1 + ( ) 24 24
67
Step 3 Determine representative values of soil elastic modulus from soil testing and Poissons ratio The
soil elastic modulus was taken as the average value between depths of Df and 15middotB (6 feet and 24 feet)
433 + 557 + 695 + 590 + 501 Es = = 555 tsf = 708 psi
5
ldquoDrained sandsiltrdquo gives a ν = 020
Step 4 Calculate the net cone tip resistance
qtnet = qt minus σvo = 1250 psi minus 189 psi = 12311 psi
Step 5 Calculate the bearing capacity of the sand
In drained sandssilts the bearing capacity is taken as the stress when (sB) = 011 (or 11 foundation width)
minus0345 minus0345 05 s L 600 qmax = hs ∙ qtnet ∙ ( ) ∙ ( ) = 078 ∙ (9795) ∙ (011)05 ∙ ( )
B B 144
= 193 psi = 27860 psf qmax
Assuming a factor of safety (FS) of 3
qmax = 645 psi = 9287 psf
3
Step 6 Calculate settlement
68
119902119898119886119909 0345 2
1 frasl119865119878 L 119904 = 119861 ∙ [ ∙ ∙ ( ) ]
ℎ119904 119902119905119899119890119905 B
2 0345 1 645 psi 600 inches s = 144 inches ∙ [ ∙ ∙ ( ) ] = 18 inches
078 12311 psi 144 inches
Step 7 Determine if q gt qmax
q = 8000 psf lt 9287psf = qmax3
69
CHAPTER 4 DIRECT CPT METHOD FOR DEEP FOUNDATIONS
41 INTRODUCTION
The axial compression capacity (Qtotal) for a single pile includes a side component (Qside) end bearing
component (Qbase) and pile weight (Wpile) as shown in Figure 30 Since piles commonly push through
several layers a summation of the unit side frictions acting on the pile segments must be considered
over the length of the pile While the examples shown in this Guide resemble hand calculations
computer software is more efficient Programming the procedure is possible and represents a practical
method of designing deep foundations However commercial software is frequently available and is
MnDOTrsquos most common method of designing deep foundations
There are upwards of 40 different direct CPT methods that have been developed over the past five
decades to determine a piles axial compression capacity Many of the earliest methods relied on hand-
recorded information where qc data from mechanical CPTs would be collected at 20 cm intervals
whereas the direct CPT method uses scaled penetrometer readings via specified algorithms to obtain
the pile unit side friction and unit end bearing The method that will be used for deep foundation design
is the Modified UniCone method which uses all three readings of the electronic piezocone (qt fs and u2)
while addressing a variety of pile foundation types
Figure 30 Direct CPT evaluation of axial pile capacity
70
The modified UniCone Method is based upon a total of 330 pile load tests (three times the original
Unicone database) that were associated with SCPTu data Originally the UniCone method provided
approximate soil classification in five groups via a chart of qE vs fs (Figure 31) where qE = qt -u2 Later
using the modified approach with a larger data set provided soil sub classifications as shown in Figure
32 This new 9-zone normalized soil behavior type is determined using CPT data in combination with the
CPT Material Index
Figure 31 UniCone Method soil behavior type using CPT (Mayne 2017)
Figure 32 Modified UniCone Method soil behavior type using CPT (Mayne 2017)
71
42 MODIFIED UNICONE METHOD
In order to determine the axial pile capacity using the modified method the first step requires the
determination of geoparameters as shown in Direct CPT Method for Soil Characterization Once the
soil unit weight and CPT Material index are determined for each soil layer then the effective cone
resistance pile unit side friction (fp) and pile end bearing resistance (qb) can be determined
421 Step 1
Work through the steps provided in CPT Method for Soil Characterization until each soil layer is
defined by its CPT material index and soil behavior type using Figure 6 After these steps have been
completed continue to Step 2 to determine qE qb and fp
422 Step 2
Once qt is determined for each layer the effective cone resistance can be calculated using Equation 37
119902119864 = 119902119905 minus 1199062 37
Where (a) qE is the specific value at each elevation along the pile sides for determining fp and (b) at the bottom of the pile qE is averaged in the vicinity of the pile tip from the tip bearing elevation to about one diameter beneath the tip for determining qb
Using CPT material index and qE the pile end bearing resistance is calculated using Equation 38
38 119902119887 = 119902119864 ∙ 10(0325∙119868119888minus1218)
The pile unit side friction is obtained from qE and Ic
39 119891119901 = 119902119864 ∙ 120579119875119879 ∙ 120579119879119862 ∙ 120579119877119860119879119864 ∙ 10(0732∙119868119888minus3605)
Where θPT = coefficient for pile type (084 for bored 102 for jacked 113 for driven piles) θTC = coefficient for loading direction (111 for compression and 085 for tension) θRATE = rate coefficient applied to soils in SBT zone 1 through 7 (109 for constant rate of penetration test and 097 for maintained load tests)
72
423 Step 3
Due to piles extending through multiple layers the unit side components acting on various pile
segments would need to be summed if not using the direct CPT method Since CPT calculates data at
regular intervals of 2 cm to 5 cm along the sides of the pile the average fp in each layer can be used
directly in Equation 40 to obtain the shaft capacity
119876119904119894119889119890 = 119891119901 ∙ 119860119904 40
Where fp = average pile side friction along pile length from eqn 39 As = πmiddotdmiddotH d = pile diameter and H = length embedded below grade
The base capacity for a pile in compression loading is given by Equation 41 For piles in tension (or uplift)
Qbase can be taken as 0
= 119902119887 ∙ 119860119887 41 119876119887119886119904119890 Where
qb = end bearing resistance from eqn 38 Ab = πmiddotd24 (area of a circular pile)
424 Step 4
The final step is to calculate the axial pile capacity using Equation 42
119876119905119900119905119886119897 = 119876119904119894119889119890 + 119876119887119886119904119890 minus 119882119901119894119897119890 42
43 AXIAL PILE DISPLACEMENTS
Movement of pile foundations can be assessed using elastic continuum theory which has been
developed using finite element analyses boundary elements and analytical closed-form solutions In
the case of piles passing through several soil layers the elastic solution can be used by stacking pile
segments (each with its own stiffness) as represented by soil Youngrsquos modulus The use of software is
recommended for pile groups Several available programs such as DEFPIG GROUP and PIGLET can
handle pile groups under axial and lateralmoment loading
73
44 EXAMPLE PROBLEMS
441 Example 5 Direct CPT Methods Axial Pile Capacity
Several piles need to be placed beneath the edge of a building Use the CPT data collected for this site
shown in Figure 34 to determine the axial capacity for one of the piles Assume round steel driven piles
will be used with a diameter of 1275 inches wall thickness of 025 inches and lengths of 80 feet
Concrete will be used to fill the piles To solve for all geoparameters use the Direct CPT Method for Soil
Characterization section All sand layers can be assumed ldquodrainedrdquo with ν = 02
Figure 33 Deep foundation end bearing pile diagram
74
Figure 34 CPT data from Minnesota for example problem 5
75
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 35 Soil layers using ldquorules of thumbrdquo for pile capacity example using direct CPT method
Layer 1
From Figure 35 fs = 18 psi taken as a representative value of the layer
76
18 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1210 pcf
145 psi
Layer 2
From Figure 35 fs = 12 psi taken as a representative value of the layer
12psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 3
From Figure 35 fs = 20 psi taken as a representative value of the layer
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
CPT Material Index
Layer 1
From Figure 35 fs = 18 psi and qc = 3000 psi
qt = qc + u2 ∙ (1 minus a) = 3000 psi + 20 psi ∙ (1 minus 08) = 3004 psi
∙ 4 feet = 1219 pcf ∙ 4 feet = 4877 psf = 34 psi σvo = γt
fs 18 psi Fr() = 100 ∙ = 100 ∙ = 060
(qt minus σvo) (3004 psi minus 34 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
77
uo = 0 psi
prime σvo = σvo minus uo = 34 psi minus 0 psi = 34 psi
(qt minus σvo)σatm (3004 psi minus 34 psi )145 psi = = Qtn prime )n (σvoσatm (34 psi145 psi)n
prime σvo 34 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(060)]2
Qtn = 3593 n = 038 lt 10 Ic = 14 (ie sand)
Layer 2
From Figure 35 fs = 12 psi and qc = 500 psi
qt = qc + u2 ∙ (1 minus a) = 500 psi + 40 psi ∙ (1 minus 08) = 508 psi
= 4838 psf + γt ∙ 45 feet = 4838 psf + 1172 pcf ∙ 45 feet = 5756 psf = 400 psi σvo
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 260
(qt minus σvo) (508 psi minus 400 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
78
uo = 0 psi
prime σvo = σvo minus uo = 400 psi minus 0 psi = 400 psi
(qt minus σvo)σatm (508 psi minus 400 psi )145 psi = = Qtn prime (σvoσatm)n (400 psi145 psi)n
prime σvo 400 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(260)]2
Qtn = 117 n = 10 le 10 Ic = 29 (ie clayey silt)
Layer 3
From Figure 35 fs = 20 psi and qc = 5000 psi
qt = qc + u2 ∙ (1 minus a) = 5000 psi + 0 psi ∙ (1 minus 08) = 5000 psi
= 5756 psf + γt ∙ 6 feet = 5756 psf + 1219 pcf ∙ 6 feet = 6488 psf = 451 psi σvo
fs 20 psi Fr() = 100 ∙ = 100 ∙ = 040
(qt minus σvo) (5000 psi minus 451 psi)
79
Step 2 Iterate to solve for Ic Steps are not shown for brevity
prime σvo = σvo minus uo = 451 psi minus 0 psi = 451 psi
(qt minus σvo)σatm (5000 psi minus 451 psi )145 psi = = Qtn prime )n (σvoσatm (451 psi145 psi)n
prime σvo 451 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(040)]2
= 1801 n = 06 lt 10 = 15 (ie sand) Qtn Ic
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic (14) Qtn (359) and Fr (06) the first layer is defined as a ldquoGravelly Sandrdquo
from Figure 36
80
Figure 36 Soil layer 1 using SBT method
81
Layer 2
Based on values of Ic (29) Qtn (117) and Fr (26) the second layer is defined as a ldquoSilty Mixrdquo
from Figure 37
Figure 37 Soil layer 2 using SBT method
82
Layer 3
Based on values of Ic (15) Qtn (180) and Fr (04) the third layer is defined as a ldquoSandrdquo from
Figure 38
Figure 38 Soil layer 3 using SBT method
Effective Cone Resistance
Layer 1
83
qE = qt minus u2 = 3004 psi minus 20 psi = 2984 psi
Layer 2
qE = qt minus u2 = 508 psi minus 40 psi = 468 psi
Layer 3
qE = qt minus u2 = 5000 psi minus 0 psi = 5000 psi
Pile End Bearing Resistance
Layer 3
= qE ∙ 10(0325∙Icminus1218) qb = 5000 psi ∙ 10(0325∙15minus1218) = 9085 psi
Pile Unit Side Friction
Layer 1
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 2984 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙14minus3605) = 99 psi
Layer 2
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 468 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙29minus3605) = 211 psi
Layer 3
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 5000 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙15minus3605) = 202 psi
84
Axial Pile Capacity
The side capacity is based on pile design as shown in Figure 39
Layer 1
Qside = fp ∙ As = 99 psi ∙ (π ∙ d ∙ L) = 99 psi ∙ (π ∙ 1275 in ∙ 48 in) = 19064 lb
Layer 2
Qside = fp ∙ As = 211 psi ∙ (π ∙ d ∙ L) = 211 psi ∙ (π ∙ 1275 in ∙ 588 in) = 457122 lb
Layer 3
Qside = fp ∙ As = 202 psi ∙ (π ∙ d ∙ L) = 202 psi ∙ (π ∙ 1275 in ∙ 240 in) = 58170 lb
Figure 39 Soil layering compared to pilings
85
End Bearing Resistance in Layer 3
d2 1275 in2
Qbase = qb ∙ Ab = 9085 psi ∙ (π ∙ ) = 9085 psi ∙ (π ∙ ) = 115932 lb 4 4
Wp = (γsteel ∙ Asteel ∙ Depth) + (γsteel ∙ Aconc ∙ Depth)
Wp = (490 pcf ∙ 003 ft2 ∙ 55 ft) + (150 pcf ∙ 085 ft2 ∙ 55 ft) = 7724 lbs
Layer 3
Qtotal = ΣQside + Qblayer3 minus Wp
Qtotal = (19064 + 45713 + 58170) lbs + 115932 lbs minus 7724 lbs = 642563 lbs
= 642 kips Qtotal
Table 8 Summary of selected and calculated parameters
Layer L (in) qc
(psi)
fs
(psi)
Fr Ic Qtn qE
(psi)
qb
(psi)
Qbase
(lb)
fp
(psi)
Qside
(lb)
Qtotal
(kip)
1 48 3000 18 06 14 356 2984 NA NA 99 19064
2 588 500 12 26 29 117 468 NA NA 211 457122
3 240 5000 20 04 15 180 5000 9085 115932 202 58170
Total 115932 534355 642
86
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Eslami A Alimirzaei M Aflaki E and Molaabasi H (2017) Deltaic soil behavior classification using CPTu records Marine Georesources amp Geotechnology 35 (1) 62-79
88
Eslami A and Gholami M (2005) Bearing capacity analysis of shallow foundations from CPT data 1463-1466 Proc ICSMGE- 16 Vol 3 (Osaka) Millpress Rotterdam
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Gavin K Adekunte A and OKelly B (2009) A field investigation of vertical footing response on sand Geotechnical Engineering 162 257-267
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Hunt CE Pestana JM Bray JD and Riemer M (2002) Effect of pile driving on static and dynamic properties of soft clay Journal of Geotechnical amp Geoenvironmental Engineering 128 (1) 13-24
Jamiolkowski M Ladd CC Germaine JT and Lancellotta R (1985) New developments in field and lab testing of soils Proc ICSMFE-11 1 57-154
Jardine R Chow F Overy R and Standing J (2005) ICP design methods for driven piles in sands and clays London Thomas Telford Publishing
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89
Karlsrud K (2012) Prediction of load-displacement behavior and capacity of axially-loaded piles in clay based on analyses and interpretation of pile load test results PhD Dissertation Norwegian Univ Science amp Technology Trondheim Norway
Karlsrud K Clausen CJF and Aas PM (2005) Bearing capacity of driven piles in clay the NGI approach Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis
Karlsrud K Lunne T Kort DA and Strandvik S (2001) CPTU correlations for clays Proc 16th ICSMGE 2 683-702
Kimura T Kusakabe O and Saitoh K (1985) Geotechnical model tests of bearing capacity problems in a centrifuge Geotechnique 35 (1) 33-45
Knappett JA and Craig RF (2012) Craigs Soil Mechanics 8th Edition London Spon Press Taylor amp Francis
Kolk HJ and Velde EVD (1996) A reliable method to determine friction capacity of piles driven into clays Paper No 7993 Proc Offshore Technology Conference Houston
Kolk HJ Baaijens AE and Senders M (2005) Design criteria for pipe piles in silica sands Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis Group
Kulhawy FH (2004) On the axial behavior of drilled shafts Geosupport 2004 1-18
Kulhawy FH and Jackson CS (1989) Foundation Engineering Current Principles amp Practices 2 (GSP 22) Reston VA ASCE
Kulhawy FH and Mayne PW (1990) Manual for Estimating Soil Properties for Foundation Design EPRI Report EL-6800 Electric Power Research Institute Palo Alto 306 page Download from wwwepricom
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Ku T and Mayne PW (2015) In-situ lateral stress coefficient (K0) from shear wave velocity measurements in soils Journal of Geotechnical amp Geoenvironmental Engineering 141 (12) 101061(ASCE) GT1943-56060001354
Ladd CC and DeGroot DJ (2003) Recommended practice for soft ground site characterization Soil amp Rock America 2003 3-57
Larsson R (1997) Investigations and Load Tests in Silty Soils SGI Report R-54 Linkoumlping Swedish Geotechnical Institute
Larsson R (2001) Investigations and Load Tests in Clay Till SGI Report R-59 Linkoumlping Swedish Geotechnical Institute
Lee J and Salgado R (2005) Estimation of bearing capacity of circular footings on sands based on cone penetration tests Journal of Geotechnical amp Geoenvironmental Engineering 131 (4) 442-452
90
Lehane BM (2003) Vertically loaded shallow foundation on soft clayey silt Geotechnical Engineering 156 (1) Thomas Telford London 17-26
Lehane BM (2013) Relating foundation capacity in sands to CPT qc Geotechnical amp Geophysical Site Characterization 1 London Taylor amp Francis Group
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Low HE Lunne T Andersen KH Sjursen MA Li X and Randolph MF 2010 Estimation of intact and remoulded undrained shear strengths from penetration tests in soft clays Geotechnique 60 (11) 843-859
Lunne T Robertson PK and Powell JJM (1997) Cone Penetration Testing in Geotechnical Practice London RoutledgeBlackie Academic
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Lutenegger AJ and Adams MT (2003) Characteristic load-settlement behavior of shallow foundations Proc FONDSUP 2 381-392
Mase T and Hashiguchi K (2009) Numerical analysis of footing settlement problem by subloading surface model Soils amp Foundations 49 (2) 207-220
Mayne PW (1987) Determining preconsolidation stress and penetration pore pressures from DMT contact pressures Geotechnical Testing Journal 10(3) 146-150
Mayne PW (1991) Determination of OCR in Clays by Piezocone Tests Using Cavity Expansion and Critical State Concepts Soils and Foundations 31 (2) 65-76
Mayne PW (2007) NCHRP Synthesis 368 on Cone Penetration Test Washington DC Transportation Research Board National Academies Press wwwtrborg
Mayne PW (2009) Geoengineering Design Using the Cone Penetration Test Richmond BC Canada ConeTec Inc
Mayne PW (2014) Keynote Interpretation of geotechnical parameters from seismic piezocone tests Proceedings 3rd Intl Symp on Cone Penetration Testing 47-73
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91
Mayne PW (2017) Stress history of soils from cone penetration tests 34th Manual Rocha Lecture Soils amp Rocks Vol 40 (3) Saouml Paulo ABMS
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Meyerhof GG (1956) Penetration tests and bearing capacity of cohesionless soils Journal of Soil Mechanics amp Foundations Division 82 (SM1) 1-19
Meyerhof GG (1974) General Report Penetration testing outside Europe Proceedings ESOPT-1 Vol 21 Swedish Geotechnical Society Stockholm
Meyerhof GG (1976) Bearing capacity and settlement of pile foundations Journal of Geotechnical Engineering Division 102(3) 197-228
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92
Mlynarek Z Wierzbicki J Goglik S and Bogucki M (2014) Shear strength and deformation parameters of peat and gyttja from CPTu DMT and VST Proceedings 5th International Workshop on CPTu and DMT in soft clays and organic soils 193-210 Poznan Poland Polish Committee on Geotechnics Menard Polska Tensar Internationa
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Nowacki F Karlsrud K and Sparrevik P (1996) Comparison of recent tests on OC clay and implications for design Large Scale Pile Tests in Clay London Thomas Telford
ONeill MW (2001) Side resistance in piles and drilled shafts Journal of Geotechnical amp Geoenvironmental Engineering 127 (1) 3-16
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Poulos HG and Davis EH (1974) Elastic Solutions for Soil and Rock Mechanics New York Wiley amp Sons
Poulos HG and Davis E (1980) Pile Foundation Analysis amp Design New York John Wiley amp Sons
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Randolph MF (2003) Rankine Lecture Science and empiricism in pile foundation design Geotechnique 53 (10) 847-875
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Robertson PK (1991) Estimation of foundation settlements in sand from CPT Geotechnical Engineering Congress 2 Reston Virginia ASCE
93
Robertson PK (2009) Cone penetration testing a unified approach Canadian Geotechnical J 46 (11) 1337-1355
Robertson PK (2009a) Performance based earthquake design using the CPT Keynote Lecture International Conference on Performance-Based Design in Earthquake Geotechnical Engineering IS Tokyo Tsukuba Japan
Robertson PK (2009b) Interpretation of cone penetration tests a unified approach Canadian Geotechnical Journal 46 (11) 1337-1355
Robertson PK and Cabal KL (2007) Guide to cone penetration testing Second Edition Signal Hill California Gregg In-Situ
Robertson PK and Cabal KL (2015) Guide to Cone Penetration Testing for Geotechnical Engineering 6th Edition Signal Hill California Gregg Drilling amp Testing Inc
Schmertmann JH (1970) Static cone to compute static settlement over sand Journal of the Soil Mechanics and Foundations Division 95 (SM3) 1011-1043
Schmertmann JH (1978) Guidelines for cone penetration test performance and design Report FHWA-TS-78-209 Washington DC Federal Highway Administration
Schnaid F Wood WR Smith AKC and Jubb P (1993) Investigation of bearing capacity and settlements of soft clay deposits at Shellhaven Predictive Soil Mechanics London Thomas Telford
Schneider JA Xu X and Lehane BM (2008) Database assessment of CPT-based design methods for axial capacity of driven piles in siliceous sands J Geotechnical and Geoenvironmental Engineering 134 (9) 1227-1244
Semple D (1980) Recent Developments in the Design amp Construction of Piles London Institution of Civil Engineers
Senneset K Sandven R amp Janbu N (1989) Evaluation of soil parameters from piezocone tests In-Situ Testing of Soil Properties for Transportation Transportation Research Record 1235 24-37
Shahri AA Melehmir A and Juhlin C (2015) Soil classification analysis based on piezocone penetration test data Engineering Geology 189 32-47
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Takesue K Sasao H and Matsumoto T (1998) Correlation between ultimate pile skin friction and CPT data Geotechnical Site Characterization 2 1177ndash1182
Tand KE Funegard EG and Briaud J-L (1986) Bearing capacity of footings on clay CPT method Use of In-Situ Tests in Geotechnical Engineering GSP 6 1017-1033
Tand KE Warden PE and Funegaringrd EG (1995) Predicted-measured bearing capacity of shallow footings on sand PISCPT 2 589-594
Thomas D (1968) Deep soundings test results and the settlement of spread footings on normally consolidated sands Geotechnique 18 (2) 472-488
94
Tomlinson MJ (1957) The adhesion of piles driven into clay soils Proc 4th Intl Conf on Soil Mechanics and Foundation Engineering 2 66-71
Tomlinson MJ (1986) Foundation Design amp Construction London Longman Scientific
Tomlinson MJ (1987) Pile Design and Construction Practice London Viewpoint Publication
Trofimenkov JG (1974) Penetration testing in the USSR Proceedings ESOPT-1 1 147-154
Tumay MT Hatipkarasulu Y Marx ER and Cotton B (2013) CPTPCPT-based organic material profiling Proc 18th Intl Conf on Soil Mechanics amp Geotechnical Engineering Paris 633-636
Uzielli M and Mayne PW (2011) Statistical characterization and stochastic simulation of load-displacement behavior of shallow footings Proc Georisk 2011 Geotechnical Risk Management amp Assessment (GSP 224) 672-679
Uzielli M and Mayne PW (2012) Load-displacement uncertainty of vertically-loaded shallow footings on sands and effects on probabilistic settlement estimation Georisk Assessment and Management of Risk for Engineered Systems and GeoHazards 6 (1) 50-69
Valsson SM (2016) Detecting quick clay with CPTu Proceedings 17th Nordic Geotechnical Meeting Reykajavik Iceland Challenges in Nordic Geotechnics
Van Dijk BFJ and Kolk HJ (2011) CPT-based design method for axial capacity of offshore piles in clay Frontiers in Offshore Geotechnics II 555-560
Vesić AS (1975) Bearing capacity of shallow foundations Foundation Engineering Handbook New York Van Nostrand Reinhold Publishers
Vesic AS (1977) NCHRP Synthesis of Highway Practice 42 Design of Pile Foundations Washington DC Transportation Research Board
Yu F and Yang J (2012) Base capacity of open-ended steel pipe piles in sand J Geotechnical and Geoenvironmental Engineering 138 (9) 1116-1128
Zawrzykraj P Rydelek P and Bakowska A (2017) Geoengineering properties of Eemian peats from Radsymin (central Poland) in the light of static cone penetration and dilatometer tests Engineering Geology 226 290-300
95
APPENDIX A
List of Figures
Figure A1 Conventional methods versus direct CPT approach to shallow foundation response A-3
Figure A2 Direct bearing capacity relationship for embedded square and strip footings situated on clean
sands (after Schmertmann 1978) A-6
Figure A3 Direct CPT bearing factor Rk as function of footing width B and embedment depth from finite
element analyses (Tand et al 1995) A-6
Figure A4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio and sand
consistency (Eslaamizaad amp Robertson 1996) A-7
Figure A5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp Salgado (2005) in
terms of footing size relative density and base settlement A-8
Figure A6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and normalized
cone tip resistance (after Eslami amp Gholami 2005) A-8
Figure A7 Direct relationship between ultimate bearing stress in clays and measured cone tip resistance
(Schmertmann 1978) A-10
Figure A8 Direct relationship between ultimate foundation bearing stress in clays and net cone tip
resistance (after Tand et al 1986) A-11
Figure A9 Measured load-displacements for each of the five large footings at TAMU (Briaud amp Gibbens
1994) sand site A-12
Figure A10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU (Briaud amp
Gibbens 1994) footings A-13
Figure A11 Characteristic stress versus square root normalized displacements for TAMU footings A-15
Figure A12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sand A-16
Figure A13 Summary of sand formation factors with corresponding CPT resistances A-19
Figure A14 Normalized foundation stresses vs square root of normalized displacement for spread
footings on sand (modified after Mayne et al 2012) A-20
Figure A15 Definitions of capacity from three types of observed foundation load test responses
(modified after Kulhawy 2004) A-21
Figure A16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footings A-22
A-1
Figure A17 Measured load vs displacement curves for 3 shallow foundations on clay at Baytown Texas
(Stuedlein amp Holtz 2010) A-25
Figure A18 Characteristic stress vs square root of normalized displacement for all three foundations at
Baytown site (Stuedlein amp Holtz 2010) A-26
Figure A19 Normalized foundation stress vs square root of normalized displacements A-27
Figure A20 Applied foundation stress vs CPT-calculated stress for 67 footings A-28
Figure A21 Applied footing stress vs CPT calculated stress on logarithmic scales A-29
Figure A22 Summary graph and equations of direct CPT method for evaluating the vertical A-30
Figure A23 Foundation soil formation parameter hs versus CPT material index Ic A-32
Figure A24 Bearing capacity ratio qmaxqnet versus CPT material index Ic A-33
Figure A25 Influence factor for rectangular foundations from elastic theory solution (Mayne amp
Dasenbrock 2017) A-34
List of Tables
Table A1 Direct CPT methods for bearing capacity of footings on clean sands A-4
Table A2 Direct CPT methods for bearing capacity of footings on clays A-9
Table A3 Summary of large footings soil conditions and reference sources of database A-17
Table A4 Sources for shallow foundation performace A-34
A-2
A1 INTERPRETATION OF SOIL PARAMETERS FROM CONE
PENETRATION TESTS
A11 Introduction
Geotechnical site characterization is an important first step towards the evaluation of
subsurface conditions and determination of soil layering geomaterial classification and the
evaluation of soil engineering parameters for the analysis and design of foundations retaining
walls tunnels excavations embankments and slope stability Increasing use of the electronic
cone penetrometer in highway applications is prominent in the USA because the results are
obtained much faster and less expensive than traditional methods that rely on rotary drilling
augering sampling and laboratory testing Moreover the cone penetration test (CPT) provides
at least three independent and continuous readings on soil behavior that are digitally recorded
and fully available at the end of the sounding As such the reliable and consistent
interpretation of CPT data is important since civil engineering works will best realize the
efficiency and economy of this technology in constructed facilities for county state and
interstate projects
A111 Cone Penetration Testing
The cone penetrometer is an electronically-instrumented steel probe that is vertically pushed into the
ground at constant rate of 20 mms (08 insec) using a hydraulic system The penetrometer is outfitted
with load cells pressure transducers and inclinometers that measure at least three readings
continuously with depth (a) cone tip resistance qt (b) sleeve friction fs and (c) porewater pressure u2
With the latter reading the device is called a piezocone thus the designation CPTu is used to indicate
porewater pressure Additional sensors are available that can be used to measure inclination
resistivity pH temperature shear wave velocity and other readings if desired
Figure A1 shows a schematic rendering of the cone penetration test that is performed in the
field per ASTM D 5778 procedures The hydraulic system is nominally rated at 180 kN (20 tons)
capacity and often mounted on a truck but can also be positioned on tracked vehicles or
anchored frames Figure A2 shows a MnDOT cone truck that uses the full dead weight of the
rig for the hydraulic reaction forces which are applied at the center of the vehicle The cab is
enclosed so that soundings may proceed during inclement weather and the data acquisition
system and operator are protected
Figure A 1 Setup and procedures for co ne penetration testing (CPT)
A-4
Figure A2 CPT truck-mounted rig used by MnDOT
A representative sounding taken at the I-35 test site just northeast of Saint Paul is presented in Figure
A3 Here the separate profiles of qt fs and u2 with depth are shown in side-by-side plots In the last
graph the hydrostatic porewater pressure (uo) due to the groundwater table is indicated by the blue
dashed line
These readings are used to interpret the soil layers types and properties of the ground as
discussed in the following sections
A-5
Figure A3 Representative CPT sounding at I-35 test site northeast of St Paul MN showing profiles of
(a) cone resistance (b) sleeve friction and (c) penetration porewater pressures
A112 Soil Parameter Interpretation
A variety of soil engineering parameters can be interpreted from CPT results on the basis of
theoretical frameworks analytical models or numerical simulations otherwise by empirical
methods based on correlations and statistics of databases Figure A4 shows a conceptual
utilization of the readings from the CPT for interpretation of geostratigraphy compressibility
flow characteristics and stress-strain-strength behavior of soils
Table A1 lists a selection of geoparameters that have been addressed for these purposes The
most common or useful relationships will be discussed in subsequent sections and reference is
made to other sources (Lunne et al 1997 Mayne 2007 Robertson amp Cabal 2015) for additional
details
A-6
Figure A4 Conceptual post-processing of CPT results for geoparameter evaluation
Table A1 List of common geoparameters interpreted from CPT data
Symbol Parameter Remarks Notes
SBT Soil behavior type (SBT)
Ic CPT material index
γt Unit weight
ρt Mass Density
σvo Overburden stress
u0 Hydrostatic pressure
A-7
σvo Effective stress
Table A1 Continued
Symbol Parameter Remarks Notes
σp Preconsolidation stress σp = 033 (qnet)m where m = fctn (Ic)
(or Effective Yield Stress) where stresses are in kPa
YSR Yield stress ratio YSR = σpσvo (formerly OCR)
c Effective cohesion intercept
ϕ Effective friction angle
su Undrained shear strength
D Constrained modulus
G0 Small-strain modulus
G Shear modulus
E Youngs modulus
cvh Coefficient of consolidation
K Bulk modulus
A-8
LI
K0 Lateral stress coefficient
k Hydraulic conductivity
Liquefaction index
A12 Geostratigraphy and Soil Behavioral Type (SBT)
In routine CPT soil samples are not normally taken and thus the measured stress friction andor
porewater pressure readings are used to infer the types of soils This can be accomplished using
three basic approaches
a Rules of Thumb or approximate guidelines for a quick visual assessment
b Soil Behavioral Type (SBT) Charts that are based on either the three readings (qt fs u2) net readings
including net cone resistance (qnet = (qt-σvo) effective cone resistance (qE = qt - u2) friction ratio (Rf =
100middotfsqt) and excess porewater pressures or using normalized CPT readings such as Q F Bq or U
c Probabilistic methods where the CPT readings have been calibrated from lab tests on
recovered soil samples (eg Tumay et al 2013)
A121 Approximate Rules of Thumb
The approximate rules of thumb provide simple guidelines for soil type and suggest that sands are
identified when qt gt 5 MPa and u2 asymp u0 whereas intact clays are prevalent when qt lt 5 MPa and u2 gt u0
(Mayne et al 2002) The magnitude of porewater pressures reflects the consistency of the intact clay
such that soft (u2 asymp 2middotu0) firm (u2 asymp 4middotu0) stiff (u2 asymp 8middotu0) and hard (u2 asymp 20middotu0) However for fissured
overconsolidated clays the measured porewater pressures are less than hydrostatic in fact often
negative u2 lt 0 On land negative porewater pressure readings at the shoulder filter element of the
piezocone should be greater than -1 bar ie u2 gt -100 kPa (-147 psi) Also for clean sands the friction
ratio (FR = 100middotfsqt lt 1) while for insensitive clays (FR gt 4)
A-9
Figure A5 presents a 36-m deep piezocone sounding from the MnDOT Wakota Bridge site which crosses
the Mississippi River at I-494 (Dasenbrock 2006) The qt and u2 readings have been annotated using the
aforementioned rules of thumb indicating the predominance of sands at this site As noted there are
five interbedded clay layers evident at depths of 1 m 3-4 m 7 - 12 m 17 m and 22-27 m
Figure A5 Interpretation of soil types from CPT data taken from the Wakota Bridge on I-494 using
approximate rules of thumb
A122 Soil Behavioral Type Charts
The most common approach for identification of soil types is based on soil behavioral charts
Reviews of several chart methods are provided elsewhere (Kulhawy amp Mayne 1990 Fellenius amp
Eslami 2000 Shahri et al 2015) Current popularity favors the 9-zone classification system to
evaluate soil behavioral type (SBT) that uses normalized piezocone readings (Robertson 1990
1991 Lunne et al 1997)
A-10
(119902119905 minus 120590119907119900) (a) normalized tip resistance 119876 = A11 prime 120590119907119900
100∙119891119904 (b) normalized sleeve friction 119865() = A12
(119902119905 minus 120590119907119900)
(1199062 minus 1199060) (c) normalized porewater pressure 119861119902 = A13
(119902119905 minus 120590119907119900)
While the Q-F-Bq classification is a three-dimensional plot it is often presented as a set of two-dimensional
plots specifically Q vs F and Q vs Bq as shown in Figure A6 Data are grouped according to layers and
can be superimposed to identify their association with the nine soil behavioral types All indirect CPT soil
classification approaches should be cross-checked and verified for a particular geologic setting and local
geotechnical conditions before routine use in practice
Figure A6 Nine-zone soil behavioral type charts (a) normalized cone resistance vs normalized
sleeve friction (b) normalized cone resistance vs normalized porewater pressure parameters
(adapted after Robertson 1991 Lunne et al 1997)
The development of a CPT material index (Ic) has been found advantageous in the initial screening of soil
types and helps to organize the sounding into their respective SBT zones of similar soil response In this
case the CPT index is found from (Robertson 2009)
A-11
A2 22 )log221()log473( rtnc FQI
where Qtn = modfied stress-normalized net cone resistance given by
(119902119905 minus 120590119907119900)120590119886119905119898 119876 = A31 prime (120590119907119900120590119886119905119898)119899
where σatm = reference stress equal to atmospheric pressure (1 atm asymp 1 bar = 100 kPa asymp 1 tsf asymp 147 psi)
The exponent n is soil-type dependent n = 1 (clays) n asymp 075 (silts) and n asymp 05 (sands) If units of bars are
used then Qtn (units of bars) is simply
(119902119905 minus 120590119907119900) 119876 = 119899 A32
prime ) (120590119907119900
The operational value of exponent n is found by iteration Initially an exponent n = 1 is used to calculate
the starting value of Ic (ie Qtn = Q) and then the exponent is upgraded to
prime 120590119907119900 119899 = 0381 ∙ 119868119888 + 005 ( ) minus 015 le 10 A4 120590119886119905119898
Then the index Ic is recalculated Iteration converges quickly and generally only 3 cycles are needed to
secure the operational Ic at each depth The soil zones and associated Ic values are detailed in Figure A7
The sensitive soils of zone 1 can be screened using the following expression
A-12
Zone 1 119876119905119899 lt 12119890119909119901( minus14 ∙ 119865119903)
A5
If any soil layers are found within zone 1 (sensitive soils) then caution should be exercised as these
structured clays are prone to instability collapses and difficulties in construction and performance
Another indicator of zone 1 sensitive soils is when the normalized porewater parameter Bq gt 08 When
these criteria are evident the geoengineer should consult with senior geotechnical personnel or the chief
engineer for guidance
Figure A7 Delineation of soil behavioral type zones using CPT material index Ic
A-13
Stiff overconsolidated clayey sands and sandy clays soils of zone 8 (15 lt Fr lt 45) and zone 9 (Fr gt
45) can be identified from the following criterion
Zones 8 and 9 0020)1(00030)1(0050
12
rr
tnFF
Q A6
The remaining soil types are identified by the CPT material index Zone 2 (organic clayey soils Ic ge 360)
Zone 3 (clays to silty clays 295 le Ic lt 360) Zone 4 (silt mixtures 260 le Ic lt 295) Zone 5 (sand mixtures
205 le Ic lt 260) Zone 6 (clean sands 131 le Ic lt 205) and Zone 7 (gravelly to dense sands Ic le 131)
The red dashed line at Ic = 260 represents an approximate boundary separating drained (Ic lt 260) from
undrained behavior (Ic gt 260)
Once the specific zone has been assigned to a layer a visual representation can be made to show
either the zone number or colorization so that the predominant layers and soil types can be
realized
A123 Probabilistic Soil Types from CPT
The inference of soil type has been addressed using probabilistic methods as discussed by Abu-
Farshakh et al (2008) and Tumay et al (2013) Here the CPT readings can be post-processed to
provide the percentage of sand silt and clay components with depth The output is consistent
and compatible with the current MnDOT soil classification chart (Figure A8) which is similar in
content to that used by the USDA
A-14
Figure A8 Chart for soil classification system by MnDOT
A13 Identifying Soft Organic Soils
The identification of soil type using cone penetration tests (CPT) is usually done on the basis of empirical
soil behavioral type (SBT) charts that use the measured readings (qt fs andor u2) While there at least
25 different sets of charts available (ie Hegazy 1998 Eslami et al 2017 Valsson 2016) one of the most
widely used and popular is the 9-zone SBT system that uses three normalized cone readings Qtn Fr and
Bq (Robertson amp Cabal 2007 Lunne et al 1997) The system is denoted as SBTn and plotted on charts in
terms of Q vs F and Q vs Bq as shown in Figure A9
A-15
1
10
100
1000
01 1 10
Norm
aliz
ed
Con
e T
ip R
esis
tan
ce
Q
Normalized Friction Fr = 100 fst(qt - svo) ()
9-Zone Soil Behavioral Type Chart
Gravelly Sands(zone 7)
Sands
(zone 6)
Sandy Mixtures(zone 5)
Silt Mix(zone 4)
Clays
(zone 3)
Organic Soils(zone 2)
Sensitive Clays
and Silts(zone 1)
Very stiff OC clayto silt (zone 9)
Very stiffOC sandto clayey
sand(zone 8)
1
10
100
1000
-06 -04 -02 00 02 04 06 08 10 12 14
No
rmal
ized
Co
ne
Tip
Res
ista
nce
Q
Normalized Porewater Bq = Du2(qt-svo)
Clays(zone 3)
Sensitive(zone 1)Organic
(zone 2)
Gravelly Sands(zone 7)
Sands(zone 6)
Figure A9 CPT charts for SBTn system (a) Q versus F (b) Q versus Bq
Of particular concern in geotechnical site characterization is the proper identification of soft
organic soils such as organic clays and silts (OL and OH) and peats (Pt) These soils often cause
difficulties and problems during construction and performance of civil engineering works because
of bio-degradation long-term creep excessive settlements andor other issues
The SBTn system has a category labelled Zone 2 recognizes organic soils However several recent
studies have noted that data from CPTu soundings do not always fall within the bounds
delineated using Zone 2 even though laboratory tests and field inspections clearly note the
presence of organic soils Lab tests to classify organic soils includes loss on ignition (LOI gt 5)
and two pairs of Atterberg limits testings per ASTM standards as well as the visual-manual
identification of strong organic odor and smell and dark color notably black and dark gray
Results by Mlynarek et al (2014) show that organic soils in Poland are not adequately identified by the
Q-F chart as seen in Figure A10 Moreover studies by Zawrzykraj et al (2017) found that neither the Q-
A-16
F and Q-Bq plots worked well to recognize organic soils and peats in Poland Nejaim et al (2016) found
that Brazilian soft organic clays were misclassified as Zone 3 (clays) rather than Zone 2 (organic clays)
Similarly a recent study on CPTu data from organic clays located at 24 sites in the USA Sweden Mexico
Brazil and Australia have been compiled (Agaiby 2018) These data are shown to generally avoid falling
with the Zone 2 boundaries in either Q-F or Q-Bq charts thus are not recognized during these soundings
as indicated by Figure A11 The exception in this case is the Mexico City clay which is correctly identified
however the other 23 sites are generally not properly recognized and categorized Additional findings by
Missiaen et al (2015) found that the CPTu results in Belgian soils also did not identify organic clays and
peats in the non-normalized version of these charts that uses Qt versus friction ratio (Rf = 100middotfsqt) as
detailed by Robertson amp Cabal (2007)
Figure A10 Soil behavioral chart with superimposed CPTu data from Polish organic soils
(from Mlynarek et al 2014)
A-17
Figure A11 Paired SBTn charts with CPTu data from 24 organic clay sites indicating non-compliance
with Zone 2 (organic soils)
(compiled by Agaiby 2018)
A14 Simplified Yield Stress Evaluation in Clays by CPTu
From the development of a hybrid formulation of spherical cavity expansion (SCE) theory with critical-
state soil mechanics (CSSM) the effective preconsolidation stress or yield stress (σp) of clays can be
expressed in terms of three piezocone parameters (a) net cone tip resistance qnet = qt - σvo (b)
measured excess porewater pressure Δu2 = u2 - u0 and (c) effective cone resistance qE = qt - u2 Details
on the SCE-CSSM solutions are given elsewhere (Mayne 1991 Mayne 2017 Agaiby amp Mayne 2018)
A-18
For normal and well-behaved clays which include inorganic fine-grained soils such as clays and silts
of low sensitivity a set of simple relationships can be derived By adopting characteristic default values
for effective friction angle ϕ = 30deg and rigidity index (IR = 100) linear expressions for the effective
preconsolidation stress are obtained
prime 120590119901 = A7 033 ∙ 119902119899119890119905
prime 120590119901 = 053 ∙ ∆1199062 A8
prime 120590119901 = 060 ∙ 119902119864 A9
A141 Application to soft Chicago clay
An example of a common geomaterial in this category includes the infamous soft Chicago clays
that were deposited in a glacio-lacustrine environment (Cho amp Finno 2010) The soft clay has a
mean water content of 20 liquid limit of 38 and plasticity index PI= 12 Results of CPTu tests
conducted at the national geotechnical experimentation site (NGES) at Northwestern University
are shown in Figure A12 (Ouyang amp Mayne 2017) As seen in the profile a thin soft clay resides
between depths of 9 to 105 m with a thicker soft clay layer in the range of 12 to 18 m
A-19
0
2
4
6
8
10
12
14
16
18
20
00 02 04 06 08 10 12D
ep
th (
m)
CPTu qt and u2 (kPa)
sand (SP)
Blodgett
Park Ridge
Deerfieldsoft clay
ClayLayerof Study
Northwestern UnivNational Test Site
Figure A12 Results of CPTu sounding at the NGES at Northwestern University Illinois
Post-processing of the CPTu data using Equations A7 A8 and A9 show good agreement in evaluating the
profile of effective yield stress in this clay layer compared with results from one-dimensional
consolidation tests made on undisturbed samples taken at the site
A-20
10
11
12
13
14
15
16
17
18
19
20
0 100 200 300 400 500 600 700 800 900 1000
De
pth
(m
)Effective Yield Stress sp (kPa)
Consols
033 qnet
053 Du2
060 qE
svo
Figure A13 Evaluation of effective yield stresses at NWU using consolidation tests and CPTu
A142 Application to soft Bay Mud California
Another example of a well-behaved fine-grained soil is the San Francisco Bay Mud which is a soft clay
deposit in northern California Results of a representative CPTu are shown in Figure A14 and indicate the
soil profile at a site in eastern San Francisco (Hunt et al 2002) For this site index values include 70 lt
wn lt 90 60 lt LL lt 80 and 35 lt PI lt 45 Post-processing the CPTu data using Equations A7 A8 and
A9 show good agreement with each other as well as with the results of consolidation tests as seen in
A-21
Figure A15 The consolidation tests were performed using a CRS device as reported by Pestana et al
(2002)
San Francisco Bay Mud (Pestana et al 2002)
0
5
10
15
20
25
30
35
40
0 1000 2000 3000
De
pth
(m
ete
rs)
Cone Resistance qt (kPa)
0
5
10
15
20
25
30
35
40
0 10 20 30 40 50 60
Sleeve Friction fs (kPa)
0
5
10
15
20
25
30
35
40
0 500 1000 1500
Porewater u2 (kPa)
Miscellaneous Fill
Young Bay Mud
Young Bay Mud
Clayey Sand
Miscellaneous Fill
Young Bay Mud
Young Bay Mud
Figure A14 Representative CPTu in San Francisco Bay Mud (data from Hunt et al 2002)
A-22
San Francisco Bay Mud (Pestana et al 2002)
0
5
10
15
20
25
30
0 100 200 300 400 500 600
De
pth
(m
ete
rs)
Preconsolidation Stress sp (kPa)
svo
033 qnet
053 Du2
060 qE
CRS
svo
Figure A15 Evaluation of effective yield stresses in Bay Mud from consolidation tests and CPTu
A15 CPTu Screening of Soft Organic Clays
For soft organic clays the Equations A7 A8 and A8 will not apply thus can serve as a means to
screen such geomaterials from the soil profile Two examples of CPTu soundings in soft organic
clays are presented to illustrate the approach Additional case records and documentation of
CPTu data in organic clays are given in Agaiby (2018)
A151 Soft organic alluvial soils in Washington DC
Results of in-situ tests including CPTu soundings in soft organic clayey silts along the Potomac River at the
Anacostia Naval Air Station are presented by Mayne (1987) Here the soft soils are categorized as organic
clayey silts (OH) per the Unified Soils Classification Systems (USCS) with mean values (and plus and minus
A-23
one standard deviations) of wn = 68 plusmn 16 LL = 83 plusmn 25 and PI = 37 plusmn 17 A representative
piezocone sounding is depicted in Figure A16 For the post-processing of CPTu data the use of the Q-F
and Q-Bq graphs do not find any points within the Zone 2 category for organic soils When the data are
evaluated to determine the profile of effective yield stress the Equations A7 A8 and A9 do not agree as
shown by Figure A17
0
5
10
15
20
25
0 2000 4000 6000
Dep
th (
met
ers)
Cone Resistance qt (kPa)
0
5
10
15
20
25
0 20 40 60 80 100
Sleeve Friction fs (kPa)
0
5
10
15
20
25
0 200 400 600
Pore Pressure u2 (kPa)
Figure A16 Representative CPTu in soft organic clay along the Potomac River DC
A-24
0
5
10
15
20
25
0 3 6 9D
epth
(m
eter
s)Soil Behaviour Type Zone
Q and Fchart
0
5
10
15
20
25
0 100 200 300 400 500 600
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
033qnet
053 Du2
06 qE
Oed
Figure A17 Post-processing of CPTu data from Anacostia-Bolling (a) SBTn zones using CPT material
index Ic (b) CPTu screening equations for yield stress
Note that the hierarchy of the results from Equations A7 A8 and A9 show that the CPTu
evaluations for preconsolidation stress in soft organic clays indicates
[A10] σp = 053 Δu2 lt 033 qnet lt 060 qE
A152 Soft organic peaty clay in Saint Paul MN
Results of CPTu tests were collected during a training exercise underneath I-35E just northeast
of Saint Paul MN in 2007 All three MnDOT CPT rigs available at the time were used to obtain
in-situ test data at the site The site is well known as having soft organic soils and samples were
obtained from three borings made at the site (Boring ID T523 T542 and T518) Boring logs
A-25
indicate the presence of highly organic silt loams with marl and spongy organic clayey silts
From 12 samples the natural water contents ranged from 30 to 191 with a mean of 112 plusmn
58
A representative piezocone sounding from the site (MnDOT CPTu ID F22Y0703C) is used for
this illustration and is presented as Figure A18 Post-processing these data using the SBTn
algorithms and CPT material index show that the soils classify primarily as Zone 3 (clays) and
Zone 4 (silty mix) thus do not identify the geomaterials properly as Zone 2 (organic soils)
0
5
10
15
0 500 1000 1500 2000
Dep
th (
met
ers)
Cone Tip Resistance
qt (kPa)
0
5
10
15
0 10 20 30 40 50 60
Sleeve Friction
fs (kPa)
0
5
10
15
-100 0 100 200 300 400
Porewater Pressure
u2 (kPa)
Figure A18 MnDOT CPTu sounding at the I-35E test site near St Paul Minnesota
A-26
1
10
100
1000
01 1 10
Norm
aliz
ed
Con
e T
ip R
esis
tan
ce
Q
Normalized Friction Fr = 100 fst(qt - svo) ()
9-Zone Soil Behavioral Type Chart
St PaulGravelly Sands(zone 7)
Sands
(zone 6)
Sandy Mixtures(zone 5)
Silt Mix(zone 4)
Clays
(zone 3)
Organic Soils(zone 2)
Sensitive Clays
and Silts(zone 1)
Very stiff OC clayto silt (zone 9)
Very stiffOC sandto clayey
sand(zone 8)
1
10
100
1000
-06 -04 -02 00 02 04 06 08 10 12 14
No
rmal
ized
Co
ne
Tip
Res
ista
nce
Q
Normalized Porewater Bq = Du2(qt-svo)
Clays(zone 3)
Sensitive(zone 1)Organic
(zone 2)
Gravelly Sands(zone 7)
Sands(zone 6)
Figure A19 Post-processing St Paul sounding using the 9-zone SBTn classification system
Using the recommended approach Figure A20 shows the CPT-evaluated yield stress profiles from
equations A7 A8 and A9 Here again the three estimates of σp do not agree but show the same
hierarchy as detailed in Equation A10
A-27
0
5
10
15
20
0 50 100 150 200 250 300 350 400
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
06 qeff svo
033qnet Consols
054 Du2
Figure A20 CPTu screening of soft organic soils at St Paul test site
A16 CPTu Evaluations of Yield Stress in Soft Organic Soils
Once the presence of soft organic clays and silts is identified using the aforementioned CPT
screening algorithms the evaluation of effective yield stress is conducted using the procedure
outlined in Chapter 2 of this MnDOT manual For soft organic soils an exponent of m = 090 is
recommended (Mayne et al 2009) Therefore the preconsolidation stress is obtained from
(units of kPa)
prime = A11 120590119901 033 ∙ (119902119899119890119905)090
The algorithm can be expressed in dimensionless form by
prime = A12 120590119901 033 ∙ (119902119905 minus 120590119907119900)119898prime ∙ (120590119886119905119898100)1minus119898prime
A-28
where σatm = reference stress (= 100 kPa = 147 psi) and m = 090 for soft organic silts and clays
The approach is applied to the two former example case studies in Figure A21 (Washington DC)
and Figure A22 (St Paul MN) respectively showing good agreement with benchmark results
taken from laboratory consolidation tests on undisturbed samples of these sites in both cases
0
5
10
15
20
25
0 100 200 300 400 500 600
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
033qnet^090
Oed
Figure A21 Profiles of yield stress from consolidation tests and CPTu method for organic clayey silts at
Anacostia-Bolling site in Washington DC
A-29
0
5
10
15
20
0 20 40 60 80 100 120 140 160
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
CPT
Consols
120590119901prime = 033 119902119899119890119905
090 in kPa
Figure A22 Profiles of yield stress from consolidation tests and CPTu method for organic clayey silts at
MnDOT test site near Saint Paul Minnesota
A17 Soil Unit Weight
As shown by Figure A23 total soil unit weight (t) can be estimated from the CPT sleeve friction resistance
(Mayne 2014) The equation can be expressed by
120574119905 = 120574119908 ∙ [122 + 015 ∙ 119897119899 (100 lowast 119891119904120590119886119905119898 + 001)] A13
where w = unit weight of water and satm = atmospheric pressure
A-30
Figure A23 Evaluation of soil unit weight from CPT sleeve friction
A18 Effective Stress Friction Angle
Sands
The effective stress friction angle () is one of the most important soil properties as it governs the strength
of geomaterials as well as affects soil-pile interface and pile side friction While an effective cohesion
intercept (c) can also be considered this is usually reserved for cemented or bonded geomaterials or
unsaturated soils and may lose its magnitude with time ageing or with prolonged environmental
changes
For clean quartz to silica sands where porewater pressures are essentially hydrostatic (Bq = 0) the
following expression has been calibrated with triaxial compression test results from undisturbed sand
samples and normalized cone resistances as presented in Figure A24 (Mayne 2007 2014)
A-31
120601prime(119889119890119892) = 176deg + 110deg 119897119900119892 (1199021199051) A14
Figure A24 Evaluation of effective stress friction angle in quartz-silica sands from CPT
Note Relationship applies to drained soil behavior when Ic lt 26 andor Bq lt 01
where qt1 is an earlier form of stress normalized cone tip resistance for sands given by
(119902119905 120590119886119905119898) 1199021199051 = A151 05 prime (120590119907119900120590119886119905119898)
A-32
In terms of units of bars this is expressed simply as (in units of bars)
119902119905 1199021199051 = 05 A152 prime ) (120590119907119900
Recently Robertson amp Cabal (2015) recommended the use of the modified form of normalized cone
resistance (Qtn) in Equation A10
120601prime (119889119890119892) = 176deg + 110deg 119897119900119892 (119876119905119899) A16
In sands qnet asymp qt since the overburden term is small relative to the cone tip resistance Figure A25 shows
that the two normalizations give very comparable results
Figure A25 Comparable values from qt1 and Qtn normalization schemes for CPTs in sands
Clays and Silts
A-33
In the case of soft to firm intact clays and silty clays the effective friction angles are determined from the
normalized cone resistance and porewater pressure parameters (Senneset et al 1989 Mayne 2016) as
shown in Figure A26 The exact solution when the angle of plastification = 0 is given as
qBQ
)tan1(tan61
1)tanexp()245(tan 2
A17
Approximate NTH Solution for from CPTu
qBQ
)tan1(tan61
1)tanexp()245(tan2
Approx asymp 295deg∙Bq0121∙[0256 + 0336∙Bq + log Q ]
1
10
100
20 25 30 35 40 45
Co
ne
Res
ista
nce
Nu
mb
er Q
(degrees)
Bq
010204060810
c = 0 = 0deg
Theory (dots)
Approximation (lines)
Theory
Figure A26 Evaluation of effective stress friction angle in clays and silts from CPTu results
Note Relationship applies to undrained soil behavior when Ic ge 26 andor Bq ge 01
which can be approximately inverted into the form (Mayne 2007)
A-34
0121 120601prime = 295deg ∙ 119861119902 ∙ [ 0256 + 0336 ∙ 119861119902 + log(119876)] A18
This algorithm is specifically applicable for the following ranges of porewater pressure parameter (01 le
Bq lt 1) and effective stress friction angles (20deg le lt 45deg)
A19 Stress History
The stress history can be characterized by an apparent yield stress or preconsolidation stress (sp) as well
as by its normalized and dimensionless form YSR = spsvo = yield stress ratio The YSR is in effect the
same as overconsolidation ratio (OCR) however now generalized to accommodate mechanisms of
preconsolidation that occur beyond just erosion glaciation and removal of overburden stresses but also
due to ageing desiccation repeated cycles of wetting-drying bonding repeated freeze-thaw cycles
groundwater changes and other factors
A-35
A generalized approach for evaluating the yield stress or preconsolidation in soils using net cone
resistance has been formulated as presented in Figure A27 (Mayne 2015) Yield stress in soils from CPT
10
100
1000
10000
10 100 1000 10000 100000
Yie
ld S
tre
ss
s
y
(kP
a)
Net Cone Resistance qt - svo (kPa)
Blessington
General Trend
sy = 033(qt-svo)m
Fissured Clays m = 11
Intact clays m = 10 Sensitive clays m = 09
Silt Mixtures m = 085Silty Sands m = 080
Clean Sands m = 072
m = 11 10 09
085
080
072
Units kPa
Fissured Clays
Figure A27 Evaluation of yield stress or preconsolidation stress in soils from CPT
The algorithm can be expressed in dimensionless form by
1minus119898prime prime 120590119886119905119898 120590119901 = 033 ∙ (119902119905 minus 120590119907119900)119898prime
( ) A19 100
where m = exponent depends on soil type m = 1 (intact clays) 085 (silts) 080 (sandy silts to silty sands)
and 072 (sands) For fissured clays the exponent m may be 11 or higher depending upon the age of the
A-36
formation and degree of jointing and discontinuities Note that fissured clays can be identified when Ic lt
26 and Bq lt 01
The exponent m has also been calibrated with CPT material index as presented in Figure A28
An algorithm to express this relationship for non-fissured soils is given by
25)652(1
2801
cIm
A20
This allows the post-processing of CPT to automatically choose the appropriate exponent m for
a layer by layer analysis
Figure A28 Yield stress exponent m in terms of CPT material index Ic
A-37
A110 Lateral Stress Coefficient
The horizontal geostatic state of stress is represented by the lateral stress coefficient K0 = shosvo
commonly referred to as the at-rest condition The magnitude of K0 for soils that have been loaded and
unloaded can be approximately estimated from
119870119900 = (1 minus 119904119894119899120601prime) ∙ 119874119862119877119904119894119899120601prime A21
Data compiled from in-situ K0 measurements using self-boring pressuremeter tests (SBPMT) total stress
cells (TSC) or push-in spade cells andor laboratory methods (instrumented consolidometers triaxials
andor suction measurements) on a variety of clays silts and sands have been compiled and reported by
Ku amp Mayne (2015) as shown in Figure A29 verifying Equation A15 as a means for evaluating K0 in soils
Figure A29 Relationship between lateral stress coefficient K0 and YSR or OCR in soils (Ku amp Mayne
2015)
A-38
A111 Undrained Shear Strength
The loading of soils can occur under conditions of being fully drained (Du = 0) partially drained or full
undrained (DVV0) where Du = excess porewater pressures (above hydrostatic) and DVV0 = volumetric
strain Further details on specific stress paths are best explained in terms of critical state soil mechanics
or CSSM (Mayne et al 2009 Holtz Kovacs amp Sheahan 2011) The prevailing drainage conditions depend
upon the rate of loading and permeability characteristics of the soil Normally in sands that are pervious
and exhibit high permeability a drained response occurs Exception to this may occur in loose sands
during fast earthquake loading resulting in soil liquefaction In clays that exhibit low permeability a fast
rate of loading will result in undrained loading at constant volume This in fact is a temporary and
transient condition often termed short term loading In the long term eventually porewater pressures
will dissipate (albeit slowly) and a drained response will prevail thus termed long term loading
The overall soil strength is controlled by the effective stress strength envelope Most commonly this is
represented by a simple linear relationship termed the Mohr-Coulomb criterion where the maximum
shear stress (max called the shear strength) is given as
= 119888prime + 120590prime ∙ 119905119886119899 120601prime A22 120591119898119886119909
where c = effective cohesion intercept s = effective normal stress and = effective stress friction angle
As a starting point values of c = 0 and = 30deg can be adopted for all soil types at least until the specific
soil type geologic formation and results from high-quality laboratory or field test data are available
Peak Undrained Shear Strength from Stress History
Using simplified critical state soil mechanics the peak undrained shear strength (max = su) can be
evaluated from the effective stress strength envelope (c = 0 effective friction angle ) and stress history
(ie OCR) in the form
prime DSS 119904119906 = (119904119894119899120601prime2) ∙ (119874119862119877)Λ ∙ 120590119907119900 A23
A-39
which corresponds to a simple shear mode Figure A30 presents the aforementioned relationship
together with data from 17 different clays tested in simple shear
Figure A30 Normalized undrained shear strength from simple shear tests on various clays
versus YSR or OCR
For the triaxial compression (TC) mode the equation would give slightly higher strengths that are
calculated from
prime TC 119904119906 = (1198721198882) ∙ (1198741198621198772)Λ ∙ 120590119907119900 A24
where Mc = (6 middot sin)(3-sin)
Peak Undrained Shear Strength from CPTu
A more direct approach to assessing undrained shear strength is via bearing capacity theory
whereby
A-40
A25 kt
votu
N
qs
s
where Nkt = bearing factor that depends on mode of shearing sensitivity of the clay degree of
overconsolidation and other factors For soft offshore clays the back figured Nkt from data collected at
14 well-documented sites (Low et al 2010) determined mean values based on mode of shearing Nkt =
119 (triaxial compression) Nkt = 136 (simple shear) and Nkt = 133 (vane)
A recent study of 51 clays that were tested by both field piezocone and laboratory CAUC triaxial tests
showed that essentially Nkt = 12 for intact clays and clayey silts of low to medium sensitivity (Mayne
Peuchen amp Baltoukas 2015) Figure A31 shows the slightly different trends for offshore versus onshore
clays For sensitive clays a lower value of Nkt = 10 would be appropriate and for fissured clays a higher
value of Nkt would be in the range of 20 to 30
A-41
A-42
Figure A31 Database from 51 clays with CPT qnet versus lab measured triaxial shear strength
(Mayne Peuchen amp Baltoukas 2015)
For soft to firm clays an alternate means to evaluate undrained shear strength is via the excess porewater
pressures ( u = u2 - u0) from the following
u
uN
us
D
D 2
A26
where N u = porewater bearing factor also dependent on the aforementioned factors The study by
Low et al (2010) found N u = 59 (triaxial compression) N u = 69 (simple shear) and N u = 71 (vane
shear)
A third alternate is found from the effective cone resistance (qE = qt - u2) whereby
kE
tu
N
uqs 2
A27
where NkE is a bearing term for this approach Mayne et al (2015) indicated a mean value of NkE = 8 for
soft-firm intact clays
Remolded Undrained Shear Strength from CPT
The remolded undrained shear strength (sur) is obtained from either field vane tests lab mini-vane shear
tests or lab fall cone devices This affords the evaluation of the clay sensitivity (St) which is defined as the
ratio of peak to remolded strengths at a given water content
119904119906(119901119890119886119896) 119878119905 = frasl A28 119904119906119903
For the CPT the sleeve friction has been noted to give values that are comparable to the
remolded undrained shear strength (Powell amp Lunne 2015) Therefore
asymp 119891119904 A29 119904119906119903
A112 Ground Stiffness and Soil Moduli
The stiffness of the ground can be represented by several geoparameters including the compressibility
indices (Cr Cc Cs) spring constants (kv) and moduli Regarding the latter there are several moduli that
are used in geotechnical engineering including shear modulus (G and Gu) Youngs modulus (E and Eu)
constrained modulus (D) bulk modulus (K) resilient modulus (MR) and subgrade reaction modulus (ks)
A-43
Soil Modulus
The definition of modulus is taken as E = DsD with Figure A32 showing a typical deviator stress (q = s1
- s3) versus axial strain (a) curve Theoretical interrelationships between the elastic moduli G E D and
K have a dependence on the Poissons ratio v The value of Poissons ratio can be taken as vu = 05 for
undrained loading (ie constant volume) while for drained loading which is accompanied by volumetric
strains a value of v = 02 may be used If we adopt the reference modulus as E then the
interrelationships with the other elastic moduli are given by
a Reference Stiffness 119864prime = drained Youngs modulus A30
119864prime
b Shear Modulus 119866prime = A31 [2(1+120584prime)]
119864prime (1minus120584prime) c Constrained Modulus 119863prime = A32
[(1+120584prime)(1minus2120584prime)]
119864prime
d Bulk Modulus 119870prime = A33 [3middot(1minus2120584prime)]
A-44
Figure A32 Representative stress-strain-strength curve for soils in triaxial compression
For the extreme case when = 0 in fact D = E When = 02 the two moduli are only 10
different D = 11middotE thus the constrained modulus and drained Youngs modulus are often
considered somewhat interchangeable
The resilient modulus (MR) applies to pavement analysis and design most commonly measured by cyclic
triaxial testing under repeated load applications In fact MR is a special case of Youngs modulus that
relates to the small strain stiffness measured in the nondestructive range but has developed permanent
plastic strains after many cycles of loading (Brown 1996 Dehler amp Labuz 2007)
The subgrade reaction modulus is actually a combined soil-structural parameter as its value
depends on the ground stiffness and the size of the loaded element The subgrade modulus is
defined as ks = q where q = applied stress and = measured deflection In terms of elasticity
solutions the deflection of a flexible circle of diameter d is given by = qmiddotdmiddot(1-2)E Therefore
119864prime
119896119904 = A34 [119889middot(1minus1205842)]
which has units of kNm3 or pcf
Table A2 summarizes several selected studies towards the approximate evaluation of these moduli
obtained directly from measured CPT data Note however that the results from in-situ penetrometer
tests generally represent a peak strength as indicated in Figure A32 Thus the measured cone tip
resistance (qt) reflects the top of the stress-strain curve either the undrained strength in clays (su) or the
effective friction angle () of sands or even a strength intermediate between these two values Thus
A-45
Source Soils Studied Reference Modulus
Findings
Kulhawy amp Mayne 1990
8 stiff OC clays Lab constrained modulus
D asymp 825 middot (qt - svo)
Mohammed et al (2000)
Resilient modulus
Abu-Farsakh 2004
7 Louisiana sites
Lab constrained modulus
D asymp 358 middot (qt - svo)
Mayne (2007b)
All soil types Constrained moduli from lab consolidation
First-order estimate D asymp middot(qt - svo)
Robertson (2009)
All soil types Constrained modulus from consolidation
D = m middot (qt - svo)
when Ic gt 22
1 m = Qtn when Qtn le 14
2 m = 14 when Qtn gt 14
when Ic lt 22 then
= 003 middot 10055 Ic + 168 m
Liu et al (2016)
16 clay sites in China
Resilient modulus MR = (146qt 053 + 1355 fs
14 + 236)244
(all in MPa)
Casey et al (2016)
8 clays Triaxial tests for undrained Youngs modulus
Eu(svc)07 = 316 - 23 middot LL
where Eu and svc (MPa)
and LL = liquid limit ()
qt is probably not the best means to obtain the slope of the stress-strain curve Instead the SCPTu that
provides the initial tangent shear modulus (Gmax) from the shear wave velocity (Vs) is a better choice
Table A2 Selected Modulus Relationships with Cone Penetration Test Measurements
A-46
Nonlinear Modulus
The small-strain shear modulus (Gmax or G0) represents the initial stiffness of all soils and rocks It is the
beginning of all stress-strain-strength curves for geomaterials and obtained from elasticity theory
119866119898119886119909 = 1198660 = 120588119905 middot 1198811199042 A35
where Vs = shear wave velocity t = tga = total soil mass density t = total soil unit weight and ga =
gravitational acceleration constant (98 ms2)
From a general viewpoint on stiffness the shear modulus G of soil is defined as the slope of shear stress
versus shear strain G = DDs for a tangent definition and by G = s as a secant definition The shear
modulus is related to its associated Youngs modulus E = 2G(1+v) Both moduli are in fact highly
nonlinear ranging from a maximum value at the small-strain stiffness (Gmax = G0 = t middotVs2 where t =
total soil mass density and Vs = shear wave velocity) to intermediate G values at medium strains (asymp 1)
to low values at peak strength As such a variety of algorithms and formulae have been developed to
represent either a partial range or the full stress-strain-strength behavior of soils over a range of
interests (eg Mayne 2005) In these formulations a variable number of input parameters may be
required in order to produce a stress-strain-strength curve
A modified hyperbolic form suggested by Fahey amp Carter (1993) has favorable attributes in that the
modulus reduction factor can be established with only a single variable This allows the initial stiffness
to the be small-strain stiffness (G0) that is reduced in terms of level of mobilized strength eg max =
qqmax = 1FS which is simply the reciprocal of the calculated factor of safety (FS) The magnitude of
secant shear modulus (G) corresponding to the particular level of loading is given by
119866 = 119872119877119865 middot 1198660 = (1198661198660) middot 1198660 A36
A-47
where the MRF = modulus reduction factor determined as
1 minus (120591 )119892 119872119877119865 = (1198661198660) = frasl A37 120591119898119886119909
and g = empirical exponent term Figure A33 shows the relationships for normalized shear stress ()
versus shear strain (s) and corresponding normalized secant shear modulus (G = s) with shear strain
over a range of exponent values 01 le g le 10 For a discussion of tangent G please see Fahey amp Carter
(1993)
Using laboratory data from resonant column-torsional shear tests and triaxial specimens with local
strain measurements for Gmax reference values modulus reduction curves for a selection of sands and
clays are presented in Figure A34 The data indicate that the exponent g falls within the range 02 lt g lt
05 for many soils tested drained and undrained A mean value of g = 03 is recommended for
preliminary site investigations and designs until additional information can be obtained
The value of max is the shear strength of the soil generally taken as either (1) drained (c = 0) max =
svo middot tan or (2) undrained where max = su = undrained shear strength as discussed earlier Thus each
shear stress () can be associated with its shear modulus (G) and the relevant shear strain is found from
s = G This allows for generation of nonlinear stress-strain-strength curves at all depths from SCPTu
data in clays sands and mixed soil types
A-48
Modified Hyperbola (Fahey amp Carter 1993 Canadian Geotech J) Coefficient Term f = 1
``
0
01
02
03
04
05
06
07
08
09
1
000001 00001 0001 001 01 1
No
rmal
ized
GG
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
0
01
02
03
04
05
06
07
08
09
1
0 01 02 03 04 05 06 07 08 09 1
No
rmali
zed
GG
ma
x
Mobilized Shear Stress max
g =1
07
05
03
02
01
0
01
02
03
04
05
06
07
08
09
1
00001 0001 001 01 1
No
rmali
zed
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
0
01
02
03
04
05
06
07
08
09
1
0 1 2 3 4 5 6 7 8 9 10
No
rmali
zed
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
Reduction Factor
GGmax = 1- (max)g
Figure A33 Normalized modulus (GGmax) and normalized shear stress (max) versus shear strain (s)
for modified hyperbolic relationship of Fahey amp Carter (1993)
A-49
0
01
02
03
04
05
06
07
08
09
1
0 01 02 03 04 05 06 07 08 09 1
Mo
du
lus
Re
du
cti
on
G
Gm
ax
or
EE
ma
x
Mobilized Strength max or qqmax
NC SLB Sand
OC SLB Sand
Hamaoka Sand (e=0625)
Hamaoka Sand
Toyoura Sand e = 067
Toyoura Sand e = 083
Ham River Sand
Ticino Sand
Kentucky Clayey Sand
Kaolin
Fujinomori Clay
Pietrafitta Clay
London Clay
Vallericca Clay
Pisa Clay
Pisa Clay
Pisa Clay
Pisa Clay
Kiyohoro Silty Clay
Thanet Clay
Exp g = 02
Exp g = 03
Exp g = 04
Exp g = 05
MRF = 1 - (qqmax)g
Figure A34 Modulus Reduction Factors (MRF) with mobilized strength of clays and sands
A113 Coefficient of Consolidation
The rate at which foundation and embankment settlements occur as well as the dissipation of excess
porewater pressures is controlled by the coefficient of consolidation (cv) The magnitude of cv is also
required in the design of vertical wick drains that can be installed in soft ground to expedite the time for
consolidation Using the results of CPTu dissipation tests which measure the rate at which the u2 readings
vary with time the in-situ profile of cv can be evaluated Often the piezocone-interpreted values of cv
are validated by comparison with results from laboratory one-dimensional consolidation tests (eg Abu-
Farsakh MY 2004) Alternatively a better method is to cross-check the values with the measured full-
scale performance of instrumented embankments that are constructed over the ground and the recorded
time-rate of consolidation and settlements can provide the best cv for that site (Abu-Farsakh MY et al
2011)
A-50
Monotonic Dissipation Tests
An illustrative dissipation curve from piezocone testing at IDT State Highway 95 at an embankment and
bridge crossing is presented in Figure A35 After the CPTu sounding was halted at a depth of 512 feet
the recorded decay of porewater pressures was observed to be monotonic with time until the dissipations
were ended at 1000 seconds During that time the u2 readings decreased from 115 psi to 47 psi At this
location the depth to the groundwater table is about 16 feet thus the calculated equilibrium water
pressure is 15 psi Commonly a characteristic time for piezo-dissipations is taken at 50 degree of
consolidation although other degrees may be adopted By evaluating the value of u2 at 50 as shown in
Figure A35 the characteristic t50 = 404 s can be obtained
0
20
40
60
80
100
120
1 10 100 1000 10000
Pore
wat
er P
ress
ure
u2
(psi
)
Time (s)
CPTU-02E-1 z = 1560 m
uo (hydrostatic)
u2 (initial) = 115 psi
u2 (final projected) = u0 = 15 psi
u2 at 50 = frac12(115+15) = 65 psi
Idaho DOT State Route 95Dissipation at 512 feet
t50 = 404 seconds
Time to reach50 consolidation
A-51
Figure A35 Piezo-dissipation at Sandpoint Bridge Crossing Idaho for depth z = 512 feet illustrating the
evaluation of characteristic time t50
It is also common to plot normalized excess porewater pressures relative to the measured initial Du2
that is obtained during penetration at the constant rate of 20 mms ie DuDui versus time Figure A36
shows a set of piezo-dissipation records for a bridge and embankment site in southern Louisiana
reported by the LTRC While the porewater pressure axis is arithmetic the time scale is often plotted on
either logarithmic or square root scales In either case the t50 is then found from the measured
dissipation curve when DuDui = 050
A-52
Figure A36 Set of monotonic piezo-dissipations taken at various depths for Courtableau Bridge
Site on Louisana State Highway 103 (Abu-Farsakh et al 2011)
Dilatory Dissipation Curves
In a number of situations a monotonic type dissipation does not occur on the onset but instead a dilatory
type curve is observed Here after stopping the push of the penetrometer the recorded porewater
pressures initially begin to rise and eventually reach a peak value then afterwards follow a phase where
the Du values decrease to hydrostatic (Figure A37) A full solution is available for both cases (Burns amp
Mayne 2002) Dilatory responses are most often associated with overconsolidated clays and silts
although occasionally are observed in fine-grained soils with low OCRs
The selection of the characteristic t50 value may found by using a square root of time plot to find the initial
u2 value by projection of the post-peak dissipatory data back to the ordinate axis as shown by the example
in Figure A38 In this example the initial u2 = 480 kPa and then the same procedures in Figure A35may
apply
A-53
Figure A37 Monotonic and dilatory porewater dissipation responses in soils
(after Burns amp Mayne 1998)
Figure A38 Method for obtaining t50 from dilatory dissipation curves
(Schneider amp Hotstream 2010)
A-54
Interpretation of Dissipation Test Data
There are a number of different solutions available for the interpretation of the coefficient of
consolidation (cv) from the piezocone dissipation curves For monotonic type response the strain path
method (SPM) developed at Oxford University (Teh amp Houlsby 1991) is well-recognized while the Georgia
Tech solution by Burns amp Mayne (2002) is based on spherical cavity expansion and critical state soil
mechanics (SCE-CSSM) and handles both monotonic and dilatory curves For both solutions a simplified
procedure can be recommended that relies on the aforementioned t50 value obtained from the measured
field data as given in Table A3 The units for cv are in cm2s as shown or equivalent
Table A3 Recommended procedures for calculating cv from t50 obtained in dissipation tests
Method Equation for coefficient of
consolidation cv
RemarksNotes Eqn
No
SPM
(Teh amp
Houlsby
1991) 50
2)(2450
t
Iac
Rc
v
t50 = measured time to reach 50
dissipation
IR = Gsu = undrained rigidity index
G = shear modulus
su = undrained shear strength
ac = penetrometer radius (ac = 178
cm for 10-cm2 cone ac = 220 for a
15-cm2 size)
A32
SCE-CSSM
(Burns amp
Mayne 2002) 50
7502 )()(0300
t
Iac Rc
v
A33
A-55
Evaluating rigidity index
Both the SPM and SCE-CSSM solutions require an estimate of the in-situ rigidity index (IR = Gsu) of the
soil If the results of SCPTU are available Krage et al (2014) have derived an expression for IR that
depends upon the small-strain shear modulus net cone tip resistance and effective overburden stress
250750
max50
8111
vonet
Rq
GI
s A38
where consistent units are input for Gmax qnet and svo The value of Gmax is obtained from B29 using the
measured shear wave velocity (Vs) and unit weight evaluated from B7
If the shear wave velocity is not measured it can be estimated from the CPT data A number of
methods have been reviewed for CALTRANS by Wair et al (2012) including one that is applicable
to sands silts and clays of low sensitivity that are inorganic and uncemented
119881 (119898119904) = [101 middot 119897119900119892(119902119905) minus 114]167 middot (100 middot 119891119904119902119905)03 A39 119904
where qt is in units of kPa (Hegazy amp Mayne 1995) An alternative relationship is given by Robertson amp
Cabal (2015)
A114 Hydraulic Conductivity
The hydraulic conductivity is a parameter that expresses the flow characteristics of the soil In
geotechnics it is also called the coefficient of permeability (k) and has units of cms or feetday
Through consolidation theory the hydraulic conductivity relates directly to the coefficient of
consolidation
A-56
A40 D
ck wv
where w = unit weight of water and D = constrained modulus
Therefore one approach to evaluating k can be from the site-specific cv obtained from the
dissipation tests using an estimate of D from one of the relationships given in Table A2
An approach developed for soft normally-consolidated soils that uses t50 directly to assess the magnitude
of k is presented in
Figure A39 (Parez and Fauriel 1988) The mean trendline through the wedges of soil type can be
approximated by
251
50 (sec)251
1)(
tscmkh
A41
A-57
Figure A39 Hydraulic conductivity versus dissipation time for 50 consolidation
(Parez amp Fauriel 1988)
A-58
APPENDIX B
List of Figures
Figure B1 Conventional methods vs direct CPT approach to shallow foundation response B-3
Figure B2 Direct bearing capacity relationship for embedded square and strip footings situated on clean
sands (after Schmertmann 1978) B-6
Figure B3 Direct CPT bearing factor Rk as function of footing width B and embedment depth from finite
element analyses (Tand et al 1995) B-6
Figure B4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio and sand
consistency (Eslaamizaad amp Robertson 1996) B-7
Figure B5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp Salgado (2005) in
terms of footing size relative density and base settlement B-8
Figure B6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and normalized
cone tip resistance (after Eslami amp Gholami 2005) B-8
Figure B7 Direct relationship between ultimate bearing stress in clays and measured cone tip resistance
(Schmertmann 1978) B-10
Figure B8 Direct relationship between ultimate foundation bearing stress in clays and net cone tip
resistance (after Tand et al 1986) B-11
Figure B9 Measured load-displacements for each of the five large footings at TAMU (Briaud amp Gibbens
1994) sand site B-12
Figure B10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU (Briaud amp
Gibbens 1994) footings B-13
Figure B11 Characteristic stress versus square root normalized displacements for TAMU footingsB-15
Figure B12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sandB-16
Figure B13 Summary of sand formation factors with corresponding CPT resistancesB-19
Figure B14 Normalized foundation stresses vs square root of normalized displacement for spread
footings on sand (modified after Mayne et al 2012) B-20
Figure B15 Definitions of capacity from three types of observed foundation load test responses
(modified after Kulhawy 2004) B-21
Figure B16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footingsB-22
B-1
Figure B17 Measured load vs displacement curves for 3 shallow foundations on clay at Baytown Texas
(Stuedlein amp Holtz 2010)B-25
Figure B18 Characteristic stress vs square root of normalized displacement for all three foundations at
Baytown site (Stuedlein amp Holtz 2010) B-26
Figure B19 Normalized foundation stress vs square root of normalized displacements B-27
Figure B20 Applied foundation stress vs CPT-calculated stress for 67 footingsB-28
Figure B21 Applied footing stress vs CPT calculated stress on logarithmic scales B-29
Figure B22 Summary graph and equations of direct CPT method for evaluating the vertical B-30
Figure B23 Foundation soil formation parameter hs versus CPT material index Ic B-32
Figure B24 Bearing capacity ratio qmaxqnet versus CPT material index IcB-33
Figure B25 Influence factor for rectangular foundations from elastic theory solution (Mayne amp
Dasenbrock 2017) B-34
List of Tables
Table B1 Direct CPT methods for bearing capacity of footings on clean sandsB-4
Table B2 Direct CPT methods for bearing capacity of footings on clays B-9
Table B3 Summary of large footings soil conditions and reference sources of database B-17
Table B4 Sources for shallow foundation performace B-34
B-2
B1 DIRECT CPT METHOD FOR SHALLOW FOUNDATIONS
B11 Introduction
The analysis of shallow foundations on soils is classically handled as a two-step and two-part set of
calculations involving (a) bearing capacity commonly reliant on limit plasticity solutions and (b)
settlement or more appropriately termed displacements that are assessed via elastic continuum theory
The utilization of cone penetration tests (CPT) provides the necessary geotechnical data for the site-
specific information on the subsurface conditions at the project site including the geostratigraphy and
evaluation of input parameters This is still a viable approach where the CPT readings are interpreted to
give the soil unit weight (γt) effective friction angle (ϕ) undrained shear strength (su) preconsolidation
stress (σp) and elastic moduli (D and E) for analysis
An alternative approach is the use of CPT results to provide direct assessments of bearing capacity
andor settlements A comparison of these two distinctly different and alternate paths is depicted in
Figure B1
Figure B1 Conventional methods versus direct CPT approach to shallow foundation response
A number of available direct CPT approaches for shallow footings are reviewed within this report
Moreover as the entire load-displacement-capacity response of shallow foundations to loading occurs
as a continuous nonlinear phenomenon a single direct CPT solution is presented for the behavior of
B-3
footings on sands (drained) and clays (undrained) The methods discussed herein are specifically to
address the general case of vertical loading of shallow foundations Additional more complex situations
that consider load eccentricity moments inclined forces sloping ground and other facets may be
handled using well-established procedures that are discussed elsewhere (eg Vesić 1975 Kulhawy et
al 1983)
This report focuses on foundations on granular soils andor soils exhibiting drained behavior since
Federal Highway Administration (FHWA) studies have concluded that less than 1 of shallow
foundations for highway bridges are placed on clay soils (Paikowsky et al 2010) One reason for this is
because soft-firm intact clays require a more complicated process that assesses both a short-term
analysis (undrained) as well as long-term analysis (drained) thus requiring undrained bearing capacity
undrained distortion displacements drained bearing capacity and drained primary consolidation
settlements
B12 Direct CPT Methods for Bearing Capacity of Footings on Sands
Classical bearing capacity solutions are based most often in limit plasticity theory albeit can be
formulated from limit equilibrium cavity expansion and numerical methods Because the limit
plasticity solutions are prevalent and adopt total stress analysis they are usually developed as either
fully drained (Δu = 0) or fully undrained (ΔVV = 0) where Δu equals excess porewater pressure and
ΔVV equals volumetric strain Paikowski et al (2010) provides an extensive review of various solutions
There are a number of direct CPT methods for the evaluation of the bearing capacity of footings and
shallow foundations Table B1 lists a variety of these direct CPT approaches for footings on sands In
the direct approach a one-step process is used to scale the measured cone penetrometer readings (ie
measured cone tip resistance (qc) total cone tip resistance (qt) sleeve friction (fs) andor measured
porewater pressure u2)) to obtain the ultimate bearing stress (qult) of the foundation
Table B1 Direct CPT methods for bearing capacity of footings on clean sands
Method Surface Footing Remarks and Embedment Notes
Meyerhof (1956) qult = qc (B12) middotcw qult = qc (B12)(1+DeB)cw
Note for silty sand reduce by 05
cw = 10 dry or moist sand cw = 05 submerged sand
Meyerhof (1974) qult = qc (B + D)40 with stresses in tsf
Presumably dry sands B = fdn width (feet) and D = depth (feet)
Schmertmann (1978)
NA (applies to embedded footings)
Square qult =055σatm(qcσatm)078
Strip qult =036σatm(qcσatm)078
See Figure A2
Embedment applies De gt 05(1+B) for Blt1m De gt 12 m for B gt1m
B-4
Canadian qult = Rk0middotqc Applied to FS = 3 Geotech Society where Rk0 = 03 where FS = factor of (CFEM 199 2) safety
Tand et al (1995) NA qult = Rkmiddotqc + σvo See Figure A3 where Rk = fctn(De B)
Frank and Magnan (1995)
qult = Rk0middotqc where Rk0 = fctn(sand
consistency) Loose Rk0 = 014
Medium Rk0 = 011 Dense Rk0 = 008
qult = Rk1middotqc + σvo where Rk1 = function (De B L amp
sand consistency) Factor Rk1 = Rk0[1+Rk206+04(BL) (DeB)]
and Rk2 = 035 (loose) 050 (medium) and 085 (dense)
Loose sand qc lt 5 MPa
Medium sand 8 MPa lt qc lt 15MPa
Dense sand qc gt 20 MPa
Eslaamizaad amp Robertson (1996)
qult = KΦmiddotqc See Figure A4
See surface footing equation KΦ = function (BDe shape and density)
Lee amp Salgado (2005)
qbL = βbcmiddotqc(AVG)
where qc averaged over distance B
Not addressed See Figure A5 for factor βbc = fctn(B DR
K0 and sB)
beneath the base
Eslami amp Gholami (2005
2006)
See embedded footing solution
qult = Rk1middotqc where Rk1 = function (ratio DeB
and normalized qcσvo) See Figure A6
Measured qc and qcσvo are geometric means over 2B deep
beneath footing
Robertson amp Cabal (2007)
qult = KΦmiddotqc with KΦ = 016
See surface footing equation
Briaud (2007) qult = KΦmiddotqc with KΦ = 023
Based on full-scale tests at Texas AampM
Mayne amp Illingworth
(2010) qult = 018 qc-Mean
Note qc -Mean is averaged CPT cone resistance over depth of
influence z = 15 B deep
Based on 30 footing load tests on 12
sands
Lehane (2013) qult = 016 qc-AVE Note qc-AVE is averaged CPT cone resistance within z (m) = [B (m)
]07
Based on 47 load tests
Table B1 Continued
Notes B = minimum footing width (or diameter) L = footing length De = embedment depth cw = water
table correction σvo = total overburden stress at bearing elevation and σvo = the effective vertical stress
B-5
Figure B2 Direct bearing capacity relationship for embedded square and strip footings
situated on clean sands (after Schmertmann 1978)
Figure B3 Direct CPT bearing factor Rk as function of footing width B and embedment depth
from finite element analyses (Tand et al 1995)
B-6
Figure B4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio
and sand consistency (Eslaamizaad amp Robertson 1996)
B-7
Figure B5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp
Salgado (2005) in terms of footing size relative density and base settlement
Figure B6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and
normalized cone tip resistance (after Eslami amp Gholami 2005)
B13 Direct CPT Methods for Bearing Capacity of Footings on Clays
For direct CPT evaluations of foundation bearing capacity in clays Table B2 summarizes some of the
well-documented methods that are available Generally these assume that the loading is fast enough
and the permeability of the clay is also sufficiently low such that an undrained (ie constant volume)
condition is maintained This is fine for short term loading of foundations particularly for soft to firm
intact clays However the undrained case is not permanent Given sufficient time excess porewater
pressures in the clay will eventually dissipate and hydrostatic conditions will return to equilibrium
During this dissipation phase primary consolidation will occur that results in drained settlements Also
a drained bearing capacity condition will prevail As the CPT is advanced at a constant rate of 20 mms
the recorded readings normally constitute undrained behavior from a direct measurement viewpoint
Thus direct CPT methods are normally focused at an undrained foundation response Nevertheless the
drained footing case can be addressed using CPT results via use of conventional analysis approach and
effective stress limit plasticity solutions that require piezocone penetrometers with porewater pressure
measurements
B-8
B
D)
L
B4060(3501320kc
Many of the earlier solutions were based on data from mechanical penetrometers andor older electric
cones that measure the uncorrected tip resistance (qc) It is now well-recognized that this measurement
must be updated to the corrected cone resistance (qt) because of porewater pressure effects that act on
the mechanics of the load cell and geometry of the particular penetrometer design The results are
particularly significant in clays silts and mixed soils that are soft to firm to stiff where excess porewater
pressures are recorded
Table B2 Direct CPT methods for bearing capacity of footings on clays
Method Footing Comments NotesRemarks
Meyerhof (1974) qult = αbcmiddotqc
where 025 le αbc le 050
Applicable to saturated insensitive clays under short-term loading
Cone tip resistance
measured by
mechanical CPT
Trofimenkov (1974)
Approximation
qult asymp (qc33)09
(Assuming FS = 3)
Strip footings on clays and sandy clays 06m le B le 15m 1m le zemb le 25m
Mechanical CPT
with qc and qult in
kgcm2
Schmertmann (1978)
qult = function (qc and foundation shape)
Relationship shown in Figure A7 for square and strip footings
Cone tip resistance
measured by
mechanical CPT
Tand et al (1986)
qult = Rkmiddot(qc - σvo) + σvo
Factor Rk obtained from Figure A8 accounts for surface and deep footings Separate Rk factor for intact and jointed clays
= (qc1middotqc2)05 qc
where qc1 is geometric mean from bearing elevation to 05B deeper and qc2 is geometric mean from 05B to 15B beneath foundation base
Mix of data from
mechanical CPT qc
and electric CPT qc
LCPC Method
(Frank amp Magnan 1995)
Footing a surface
qult = kc middotqc + σvo
where kc = 032
Embedment case
Note Methods above generally consider undrained bearing capacity
B-9
Figure B7 Direct relationship between ultimate bearing stress in clays and measured cone tip
resistance (Schmertmann 1978)
B-10
Figure B8 Direct relationship between ultimate foundation bearing stress in clays and net
cone tip resistance (after Tand et al 1986)
B14 Direct CPT Assessments of Settlement and Displacements of Shallow Foundations
Methods for evaluating the magnitude of foundation displacements (s) can be found using
elastic theory subgrade reaction methods spring models and numerical simulations The
classical approach to footing settlement calculations on sands is to utilize elastic theory
solutions (eg Poulos amp Davis 1974) which take on the form
119902∙119861∙119868∙(1minus1205842) 119904 = B1
119864119904
where q = applied footing stress and I = displacement influence factor from elasticity theory
The value of I depends upon foundation geometry footing rigidity embedment depth variation of soil
modulus with depth compressible layer thickness and other factors as discussed by Mayne amp Poulos
(1999) For instance for a flexible circular footing of diameter B resting on an infinitely deep soil
formation having a constant modulus Es with depth the influence factor I is equal to 1 for the center
point The results of in-situ field tests such as pressuremeter tests (PMT) standard penetration tests
(SPT) cone penetration tests (CPT) flat dilatometer tests (DMT) and other methods can be used to
ascertain the input value of soil modulus Es
Alternatively direct CPT methods have been investigated to provide a one-step assessment of
foundation displacements Selected methods for direct CPT evaluation of shallow foundation
settlements on granular soils is include DeBeers amp Martens (1957) Meyerhof (1965) DeBeer (1965)
Thomas (1968) Schmertmann (1970) Meyerhof (1974) Berardi amp Jamiolkowski amp Lancellotta (1991)
Robertson (1991) Lutenegger amp DeGroot (1995) Lehane amp Dougherty amp Schneider (2008) Gavin amp
Adekunte amp OrsquoKelly (2009) Mayne amp Illingworth (2010) and Uzielli amp Mayne (2012)
B141 Large Scale Footing Load Tests
The measured load-displacement response of shallow foundations is conducted in the field using either
stepped loading procedures or continuous rate of displacement methods Results from the well-
documented test program at the Texas AampM University (TAMU) site (Briaud amp Gibbens 1999 Briaud
2007) for five spread footings on sand are shown in Figure B9 Each footing clearly shows a nonlinear
B-11
behavior to loading over the range of testing This site is one of the national geotechnical
experimentation sites (NGES) funded by the National Science Foundation (NSF) FHWA and American
Society of Civil Engineering (ASCE)
Figure B9 Measured load-displacements for each of the five large footings at TAMU (Briaud
amp Gibbens 1994) sand site
B142 Generalized Direct CPT Method for Footing Response on Soils
The use of applied foundation stress versus normalized displacement curves (q vs sB) is generalized to
footings on sands silts intact clays and fissured clays A square root plotting of the normalized
displacements (sB) permits a single parameter characterization for each specific soil type that in turn
correlates with the net cone tip resistance The generalization is based on a statistical review of data
from large foundations (B gt 05m) involving 70 full-scale load tests with available CPT data For fine-
grained soils the data are from electronic piezocone tests where the proper correction from qc to qt has
been obtained to allow the best possible results The methodology permits an evaluation of the load-
displacement-capacity response of shallow square footings based on CPTs performed in the four types
B-12
of ground conditions For the TAMU footings the characteristic stress-normalized displacement
response is shown in Figure B10 where all five foundations can be represented by a single curve
Figure B10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU
(Briaud amp Gibbens 1994) footings
B143 Characteristic Stress-Displacement Curves
Fellenius amp Altaee (1994) recommended the concept of a characteristic stress (q) vs normalized
displacement (sB) curve for foundation response on a given soil formation later supported by Decourt
(1999) Briaud amp Gibbens (1999) Lutenegger amp Adams (2003) and Briaud (2007) In this approach the
measured load (Q) vs displacement from individual footings of various sizes that rest on the same soil
conditions collapse to a unified relationship given in Equation B2
119939119943 119956 119954119938119953119953119949119946119942119941 = 119938119943 (119913
) B2
B-13
where af and bf are empirical fitting coefficients (Decourt 1999 Uzielli amp Mayne 2011 2012)
In a review of measured load tests data from large spread footings situated on different sands it has
been suggested that Equation B2 can be reduced to a single parameter expression by use of square root
plotting where the exponent bf is equal to 05 (Mayne amp Illingworth 2010 Mayne et al 2012)
119956 119954119938119953119953119949119946119942119941 = 119955119956radic
119913 B3
where rs = fitted soil parameter from best fit line regression In the case of sands the coefficient rs
represents the supporting soil conditions including particle size relative density and fines content as
well as geologic origin aging and overconsolidation effects
The results from the five TAMU footings are presented in this format in Figure B11 that shows the
formation factor rs = 486 MPa for this sand deposit This single coefficient can be used to express the
nonlinear load versus settlement of any size footing on this sand site Moreover one criterion from
Eurocode for bearing capacity that is used by the European community identifies that stress (or load)
which corresponds to (sB) = 10 That is when the displacement equals ten percent of the footing
width then the capacity has been reached Therefore adopting this criterion for footings on sands
the ultimate bearing stress (qult) for footings on sand can be taken when (sB) equals 010 or when
square root of (sB) equals 0316 In that case this gives qult equal to 0316 middot rs For the TAMU site this
gives qult equal to 0316 times (486) which equals 153 MPa as illustrated by Figure B12
B-14
Figure B11 Characteristic stress versus square root normalized displacements for TAMU
footings
B-15
Figure B12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sand
B15 Footing Database
A database of spread footings and large plates was assembled which considers only full-size shallow
foundations (05 le B le 6 m) that rest on sands silts and clays Sites for the foundations were also subjected to CPT The inclusion solely of large footings are significant because results from small-scale
model footings exhibit scale effects In fact experimental centrifuge work and numerical simulation
studies have shown that bearing capacity factors are size-dependent especially for factor Nγ decreasing
with footing width B (Kimura et al 1985 Cerato amp Lutenegger 2007 Mase amp Hashiguchi 2009) A
direct approach based on full-size footings provides a more reliable means for foundation evaluation
The compiled database is given in Table B3 with a total of 70 large footings and plates These include a
listing of 34 foundations on 13 sands 11 footings on 4 silt deposits 13 footings or large plate load tests
on 6 intact clays and 12 foundations on 5 fissured clays Most of the footings were square or nearly
square (80) while the remaining were circular The largest foundation was a mat 5 m by 14 m in plan
loaded by stacking Kentledge concrete blocks to cause a bearing failure in the underlying soft clay For
sands the largest footing consisted of the 6-m square concrete pad In terms of equivalent square
footings mean footing sizes were 155 m (sands) 114 m (silts) and 126 m (clays)
B-16
Additional details on the individual load tests and site conditions are given elsewhere (Mayne 2009
Mayne amp Illingworth 2010 Uzielli amp Mayne 2011 2012 Mayne et al 2012 Mayne amp Woeller 2014)
Table B3 Summary of large footings soil conditions and reference sources of database
Sand Site Location Soil Conditions Footings Numbers
Shapes and Sizes
ReferencesSource
College Station
Texas Pleistocene sand 5 Square 1 15 25 33 m Briaud amp Gibbens (1999)
Kolbyttemon Sweden Glaciofluvial sand 4 Rect B = 06 12 17 24 m
Bergdahl et al (1985 1986)
Fittija Sweden Glaciofluvial sand 3 Rect B = 06 17 24 m Bergdahl et al (1984 1985)
Alvin West Texas Alluvial sand 2 Circular D = 235 m Tand et al (1994)
Alvin East Texas Alluvial sand 2 Circular D = 22 m Tand et al (1994)
Perth Australia Silceous dune sand 4 Square B = 05 and 10 m
Lehane (2008)
Grabo T1C Sweden Compacted sand fill
1 Square B = 046 m Long (1993)
Grabo T2C Sweden Compacted sand fill
1 Square B = 063 m Long (1993)
Grabo T3C Sweden Compacted sand fill
1 Square B = 080 m Long (1993)
Labenne France Dune sand 6 Square B = 07 and 10 m
Amar et al (1998)
Green Cove Florida Brown silty sand 1 Circular D = 182 m Anderson et al (2006)
Durbin South Africa
White fine sand 1 Square B = 609 m Kantley (1965)
Porto Portugal Residual clayey sands
1 Circular D = 12 m and 1 plate
Viana da Fonseca (2003)
Jossigny France Soft clayey silt 2 Square B = 1 m Amar et al (1998)
Tornhill Sweden Glacial Baltic till 3 Square B = 05 1 2 m Larsson (2001)
Vagverkel Sweden Stiff medium silt 3 Square B = 05 1 2 m Larsson (1997)
Vattahammar Sweden Brn-Gry layered silt 3 Square B = 05 1 2 m Larsson (1997)
B-17
Table B3 Continued
Clay Site Location Soil Conditions Footings Numbers Shapes and Sizes
ReferencesSource
Baytown Texas Fissured Beaumont 2 plates with d = 076 m Stuedlein amp Holtz (2008)
Belfast Ireland Soft clay silt ldquosleechrdquo
1 Square Pad B = 20 m Lehane (Geot Engr 2003)
Bothkennar Scotland Soft silty clay 2 Square Pads B = 22 24 m
Jardine et al (1995)
Bangkok Thailand Soft to stiff clay 4 Square 067 075 090 10 m
Brand et al (1972)
Haga Norway Stiff OC clay 2 Square Footings B = 10 m
Andersen amp Stenhammar (1982)
Rio Grande Brazil Sandy residual clay 3 Square 04 07 and 10 m
Consoli et al (1998)
Shellhaven England Soft estuarine clay Rectangular 5 m by 14 m Schnaid et al (1993)
Texas City A Texas Coast Fissured Beaumont 3 circular plates d = 058 m
Tand et al (1986)
Texas City B1 Texas Coast Fissured Beaumont 2 circular plates d = 058 m
Tand et al (1986)
Texas City B2 Texas Coast Fissured Beaumont 1 circular plate d = 058 m
Tand et al (1986)
Alvin Texas Texas Coast Fissured Beaumont 3 circular plates d = 058 m
Tand et al (1986)
For the series of footings on sands Figure B13 shows the summary of measure formation factors (rs) for
the 34 foundation load tests Also indicated on the graph are the corresponding mean qc values of the
sands averaged over 15middotB beneath the bearing elevations
Note also in sands that the qt and net resistance (qnet) are all very close to the measured qc since (a)
porewater pressure corrections for the a-net value are negligible in clean sands and (b) overburden
stresses are typically only 1 to 3 of the magnitude of qc This is especially true of shallow depths
Therefore for clean sands qnet is approximately qt which is approximately qc
B-18
Figure B13 Summary of sand formation factors with corresponding CPT resistances
As suggested by Briaud (2007) various in-situ measurements can be used to normalize the results from
cone penetration tests in this case Herein the qc from the CPT soundings in the sands are used to
normalize the footing stress axis as presented in Figure B14 It is evident that the results of the load-
displacement response of all 34 footings on 13 different sands can be captured by the characteristic
stress versus normalized displacement with the inclusion of the site-specific cone tip resistance In this
case the mean trend for all footings and sand sites indicates a simple expression shown in Equation B4
119954119943119952119952119957119946119951119944 119956 Clean quartz-silica sands = 120782 120787120790120787radic B4
119954119940 119913
B-19
Figure B14 Normalized foundation stresses vs square root of normalized displacement for
spread footings on sand (modified after Mayne et al 2012)
The results of measured force-displacement curves (Q vs s) from footing load tests are nonlinear and
thus the definition of capacity must be addressed Kulhawy (2004) identified three main curve types
that occur during load testing depicted in Figure B15 The most common response (Type A) shows no
clear peak and is observed in sands and silts as well as slow loading of insensitive clays and fissured
clays In this case ldquocapacityrdquo can be defined by the Eurocode corresponding to the load (or stress) when
sB=10 (Amar et al 1998) For foundations situated on many clays a plateau capacity is reached as
depicted as Type B response In rare cases where sensitive clays or structured soils are encountered a
Type C relationship occurs whereby the load-displacement curve reaches a peak followed by strain
softening In the one case observed in this study for Type C loading (Haga clay) the corresponding
ldquopeak capacityrdquo was reached at a pseudo-strain of sB = 4 as shown in Figure B16 followed by some
strain softening In the cases of soft clays at Belfast and Bothkennar a plunging type (plateau failure)
was achieved at around sB=7
B-20
Figure B15 Definitions of capacity from three types of observed foundation load test
responses (modified after Kulhawy 2004)
B-21
Figure B16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footings
In the case of fine-grained soils that develop excess porewater pressure during cone penetration the
use of the net total cone resistance (qtnet) which equals the cone tip resistance minus the total stress
must be considered (Lunne et al 1997) As noted earlier the correction of CPT data in sands is minimal
because of the low value of penetration porewater pressures and the fact that total overburden stress is
small relative to the cone resistance especially for shallow soundings Thus the use of qtnet may be
substituted for qc into the trend of Figure B1
B16 Undrained and Drained Capacity
The footings on sands typically show drained response with no excess porewater pressures developed
during loading For the foundations situated on silts analyses by Larsson (1997 2001) concluded that
these too show drained conditions For shallow foundations on clays the entire range of drainage cases
are possible ranging from undrained to partially-drained to fully-drained depending upon the applied
rate of loading relative to the permeability of the geomaterial If sufficient data were available for each
of the clay sites then the recent evaluation of drainage condition could be assessed using normalized
velocity criteria (Chung et al 2006) However this was not possible with the current data set Instead
B-22
the test loadings of footings on clays have essentially been assumed to primarily occur under undrained
conditions following recommendations by Tand et al (1986) For this case a limiting bearing stress can
be calculated from bearing capacity analyses and related directly to the CPT readings
From classical bearing capacity calculations using limit plasticity solutions the ultimate foundation stress
is shown in Equation B5
= 119925119940 ∙ 119956119958 B5 119954119958119949119957
where Nc equals the bearing factor for constant volume (514 for strip and 614 for square or circular
plan) and su equals the undrained shear strength of the clay For the case of soft to firm clays of low to
medium sensitivity the operational strength can be related to the preconsolidation stress (Mesri 1975
Jamiolkowski et al 1985) as shown in Equation B6
prime 119956119958 = 120782 120784120784120648119953 B6
Finally the effective preconsolidation stress has been linked directly to the net cone resistance (Chen amp
Mayne 1996 Demers amp Leroueil 2002) as shown in Equation B7
prime 120648119953 = 120782 120785120785(119954119957 minus 120648119959119952) B7
Combining Equations (B5) (B6) and (B7) results in direct expressions for undrained bearing capacity on
intact clays from CPT results as shown in Equations B8 and B8-2
Strip footing 119954119958119949119957 = 120782 120785120789120785(119954119957 minus 120648119959119952) B8
B-23
Squarecircle 119902119906119897119905 = 0445(119902119905 minus 120590119907119900) B8-2
Equation (B8-2) is in excellent agreement with Tand et al (1986) database study for the case of intact
clays and footings with no embedment Their study also provided a lower relationship for foundations
situated on jointed clays specifically recommending that the bearing stress not exceed 030 qtnet
Review of the available data in Figure 3 shows this to be rather conservative and a more realistic value
may be taken at 040 qtnet for fissured clays
B161 Normalized Undrained Footing Response
The characteristic stress vs normalized displacement curves for footings on clays exhibiting undrained
conditions can be verified using a relatively recent case study on Beaumont clay by Stuedlein amp Holtz
(2010) The Baytown Texas site was investigated using an exploration program of soil borings
piezocone soundings and laboratory testing The field testing included a large square footing (B = 276
m) and two small circular plates (D = 076 m) The measured load-displacement responses for all three
foundations are presented in Figure B17 Of course the large footing carried considerably higher loads
than the two small plates on the same soil deposit Yet using the aforementioned characteristic stress
(q) vs square root of normalized displacement (sB) essentially gave a unique relationship for all 3
foundations as presented in Figure B18 In this case utilizing the form of Equation B2 the
characteristic coefficient rs is 177 MPa for this deltaic clay formation
B-24
Figure B17 Measured load vs displacement curves for 3 shallow foundations on clay at
Baytown Texas (Stuedlein amp Holtz 2010)
B-25
Figure B18 Characteristic stress vs square root of normalized displacement for all three
foundations at Baytown site (Stuedlein amp Holtz 2010)
B17 General Direct CPT Method
In a manner analogous to the normalization of applied foundation stresses shown in Figure B14 for
footings on sands a generalized direct CPT method for shallow foundations on different soils can be
made in Equation B9
120782120787 119956 119954 = 119945119956 ∙ 119954119957119951119942119957 ∙ (119913
) lt 119954119940119938119953119938119940119946119957119962 B9
where qcapacity is equal to the foundation bearing capacity of the ground and hs is equal to the empirical
fitting term that depends on soil type Specifically hs equals 058 for sands 112 for silts 147 for
fissured fine-grained soils and 270 for clays A summary of the normalized stress versus the normalized
displacement curves for these four soil types is presented in Figure B19 where the data fitting to obtain
the hs parameter on clays are limited to (sB) lt 004 corresponding to undrained loading The data on
B-26
silts and sands are considered fully drained whereas the fissured clay subset may be partially drained to
undrained The statistical measures for obtaining the fitted parameter hs are quite good as shown in
Figure B20 with the coefficient of determinations (r2) values of 0947 for sands 088 for silts 0935 for
fissured and 0925 for intact clays Additional statistical evaluations on the database are provided by
Uzielli amp Mayne (2011)
Figure B19 Normalized foundation stress vs square root of normalized displacements
B-27
Figure B20 Applied foundation stress vs CPT-calculated stress for 67 footings
The same data are shown on Figure B21 in a log-log plot to emphasize the early parts of the load-
displacement responses of these footings The best fit line from regression analyses shows excellent
statistical correlations As detailed earlier the undrained bearing capacity of square or circular footings
on clays can be taken as approximately 040 times qtnet For silts and sands that experience drained
loading with no excess porewater pressures being developed and no clear peak value for capacity the
(sB) =10 criterion can be used In this case Equation 7 can be used directly to obtain stress q equal to
qcapacity when (sB)05 is 0316
An alternate approach is presented in Figure B22 summarizing the direct CPT method for estimating the
load-displacement-capacity of shallow square footings on four soil types The maximum values of soil
bearing capacity (qmax) can be capped at stresses corresponding to a percentage of the average
measured qtnet determined over the depth interval from the foundation bearing elevation to 15middotB
below the foundation In such an approach the foundation bearing capacity can be taken as
(qcapacityqtnet) of 02 for sands 035 for silts 040 for fissured and 045 for intact clays Using the assigned
values of the hs parameter the corresponding cut-off for the general stress-normalized displacement
relationship given by Equation 7 can be established specifically giving corresponding values of the
B-28
normalized displacements which can also be considered as pseudo-strains equal to (sB) max of 4 for
clays 7 for fissured clays 10 for silts and 12 for sands
Figure B21 Applied footing stress vs CPT calculated stress on logarithmic scales
B171 CPT Soil Material Index Ic
The soil behavioral type can be assessed indirectly by CPT using a material index (Ic) as defined by
Robertson (2009a b) shown in Equation B10
119920119940 = radic[120785 120786120789 minus 119949119952119944119928119957119951]120784 + [120783 120784120784 + 119949119952119944119917119955]120784 B10
B-29
Figure B22 Summary graph and equations of direct CPT method for evaluating the vertical
where Qtn equals the stress-normalized cone tip resistance and Fr equals the normalized sleeve friction
calculated from the cone penetrometer readings as shown in Equations B11 and B12 respectably
(119954119957minus120648119959119952)120648119938119957119950 = 119951 B11 119928119957119951 prime (120648119959119952120648119938119957119950)
119943119956 119917119955() = 120783120782120782 ∙ B12 (119954119957minus120648119959119952)
where σatm equals a reference stress equal to atmospheric pressure (σatm = 1 atm asymp 1 bar asymp 100 kPa) In
the initial evaluation the exponent n is set to n = 1 to find the soil behavioral type (SBT) based on a 9-
zonal chart as shown in Figure 23 The value of exponent n varies with soil type ranging from n = 1 in
intact clays and decreasing with increasing grain size to around approximately 075 in silts and
B-30
approximately 05 plusmn 02 in clean sands The appropriate value of exponent n is found by iteration until
conversion using the relationship (Robertson 2009b) shown in Equation B13
prime 120648119959119952 119951 = 120782 120785120790120783 ∙ 119920119940 + 120782 120782120787 ( ) minus 120782 120783120787 le 120783 120782 B13 120648119938119957119950
As the distinction that footings on intact clays subjected to fast loading are behaving primarily under
undrained conditions (ie constant volume) Robertson (2009a) suggested that this occurs when Ic gt 27
while in contrast Ic lt 25 corresponds more or less to drained loading (ie no excess porewater
pressures)
The CPT material index can be used to identify soil type and the corresponding characteristic hs value for
estimating footing load-displacement response In a number of the case studies investigated the results
of the sleeve friction readings were also available to allow the calculation of Ic with depth at these sites
Figure B23 shows the tentative relationship between the values of the hs parameter plotted versus the
corresponding CPT material index with an approximate trend given by Equation B14
120784120785 119945119956 = 120784 120790 minus 120783120787 B14
119920119940 120783+( ) 120784120786
B-31
Figure B23 Foundation soil formation parameter hs versus CPT material index Ic
Soil behavioral type from the Ic ranges given by Robertson (2009b) and are shown in Figure B23 with an
overall general agreement with the known soil classifications In summary the CPT can provide an
evaluation of the load-displacement-capacity response directly via Equation B15 which may be of
interest in mixed soil types such as silty sands sandy silts and the like
120784120785 Footing stress 119954 = 119954119951119942119957 ∙ radic(119956frasl119913) ∙ [120784 120790 minus
)120783120787] B15
120783+(119920119940frasl120784120786
As discussed earlier foundation bearing capacity may be taken by a limiting value of pseudo-strain (sB)
max or by qmax as specified as a percentage of qtnet This definition of bearing capacity can be seen in
comparison to soil type as seen in Figure B24
B-32
Figure B24 Bearing capacity ratio qmaxqnet versus CPT material index Ic
B172 Rectangular Foundations
This same elastic solution from Giroud (1968) for a CPT method for square and circular foundations were
utilized for rectangular foundations The study covered rectangular distortions (AB) ranging from 1
(square) to very long foundations with AB equal to 20 where A represents the foundation length and B
represents the foundation width (Mayne amp Dasenbrock 2017) The influence factor for rectangular
shaped foundations is given in Figure B25 and the expression in Equation B16
= (119912119913)120782120785120786120787 119920119912119913 B16
B-33
Figure B25 Influence factor for rectangular foundations from elastic theory solution (Mayne
amp Dasenbrock 2017)
Including 32 full scale load tests on sands results from an additional 98 shallow foundations have been
reported in previous studies The foundations were primarily square or circular and among these include
large spread footings with measured settlements and bearing capacity Table B4 shows a summary of
the number of footings and footing sizes used in the study The range of length to width ratios varied
from 1 lt (AB) lt 23 with an average value of (AB) equal to 238 The embedment depth to footing
width ranged from 0 lt DeB lt 222 and averaged 046
Table B4 Sources for shallow foundation performace
Source of Data No of
Footings
Mean B
(m)
Max B
(m)
Min B
(m)
Mean A
(m)
Max A
(m)
Min A
(m)
Mayne et al (2012) 32 149 609 046 149 609 046
Lehane (2011) 8 041 027 060 041 060 027
Gifford et al (1987) 17 846 1588 527 1591 3596 701
Jeyapalan amp Boehm
(1986)
13 1022 2892 400 1671 3049 640
B-34
Schmertmann
(1970)
31 945 5608 061 1288 8668 061
Papadopoulos
(1992)
29 870 3600 100 1300 7290 100
Total 130 671 5608 027 1012 8668 027
The solution for shallow foundation settlements given back in Equation B1 combined with
the expression in Equation B16 takes on the form
119954∙119913∙119920119912119913∙(120783minus120642120784) 119956 = B17
119916119956
Combining the direct CPT equation developed from the 32 full scale load tests (Equation B15)
together with the influence factor for rectangular shaped foundations becomes
minus120782120785120786120787 120784120785 119912 Footing stress 119954 = 119954119951119942119957 ∙ radic(119956frasl119913) ∙ [120784 120790 minus
)120783120787] ∙ [ ] B18
120783+(119920119940frasl120784120786 119913
B18 Conclusions
A direct CPT method for square rectangular and circular shallow footings is developed using a database
of 166 full-scale field load tests where the minimum footing size of an equivalent square width of 05 m
is established to avoid scaling issues The use of load vs displacement is generalized by characteristic
stress vs normalized displacement (sB) and further simplified by a square root plotting technique that
captures the ground response in a single parameter that is related to the qtnet Data are grouped
according to four main soil categories sands silts fissured clays and intact clays It is generally believed
that the footings on sands and silts exhibit fully drained behavior while in intact clays an undrained
response occurs under conditions of constant volume
B-35
The summary equation for evaluating the vertical stress-displacement-capacity of square rectangular
and circular footings is given by
minus120782120785120786120787 119956 119912 119954119943119952119952119957119946119951119944 = 119945119956 ∙ 119954119957119951119942119957 ∙ radic ∙ ( ) lt 119954119950119938119961 B19
119913 119913
where the parameter hs also relates directly to Ic The foundation capacity depends upon the mode of
failure as well as drainage conditions and can be taken as a fraction of qtnet where qmaxqtnet is 020 in
sands 035 in silts 040 in fissured clays and 045 in intact clays as shown in Figure 24 Alternatively the
capacity can be defined using a limiting pseudo-strain given by (sB) max of 4 in clays 7 in fissured
10 in silts and 12 in sands
B-36
APPENDIX C
List of Figures
Figure C1 Components of Axial Pile Capacity C-3
Figure C2 Selected alpha relationships as function of the undrained shear strengthC-5
Figure C3 Pile side friction coefficient β in terms of effective friction angle and overconsolidation ratio
C-11
Figure C4 Concept of Direct versus Rational Method for CPT evaluation of axial pile capacityC-14
Figure C5 Soil behavior type using CPT via original UniConeC-19
Figure C6 Soil behavior type using CPT via Modified UniCone C-19
Figure C7 Nine-zone soil behavioral soil type using normalized piezocone parameters C-21
Figure C8 Summary of Modified UniCone Method for direct CPT assessment of axial pile capacity C-22
Figure C9 Elastic continuum solution for axial pile response under compression and tension loadingC-24
List of Tables
Table C1 Alpha Methods for Axial Pile Side Friction where fp = αmiddotsu C-5
Table C2 Beta Methods for Axial Pile Side Friction where fp = βmiddotσvoC-9
Table C3 Selection of Direct CPT Methods for Axial Pile CapacityC-15
C-1
C1 EVALUATING DEEP FOUNDATION RESPONSE FROM CONE
PENETRATION TESTS
C11 Introduction
The axial response of deep foundations includes the load-displacement-capacity and axial load transfer
when driven pilings and drilled shafts are loaded in compression and uplift The use of cone penetration
testing (CPT) especially seismic piezocone tests (SCPTu) are advantageous since they provide
information on the subsurface soils and their geomaterial properties The SCPTu provides at least four
measurements on soil behavior with depth including (a) cone tip resistance qt (b) sleeve friction fs (c)
penetration porewater pressure u2 and (d) shear wave velocity Vs This offers the opportunity to
determine the geostratigraphy unit weight effective overburden stress shear strength stress state
and stiffness of the ground from a single exploratory sounding thus values in economic expediency
and reliability The axial pile capacity can be calculated based on static equilibrium of forces acting along
the sides and base of the pile foundation The displacement of the pile can be ascertained using
elasticity theory from closed-form solutions boundary elements andor finite element analyses Load
transfer occurs along the length of the pile and only a portion of axial forces are transmitted to the base
or toe or tip of the pile These too can be assessed using elasticity solutions
An alternate approach to pile capacity involves the utilization of direct CPT methods available
since the 1970s but in the past 10+ years a number of new and statistically reliable algorithms
have been developed which can be implemented for highway design and construction
C111 Axial Pile Capacity
The axial compression capacity of a single pile foundation is composed of a shaft or side component and
end-bearing component at the base as depicted in Figure 1 For a circular pile the side capacity (Qs) is
determined from the unit side friction (fp) acting along the surface area of the shaft which is As = πmiddotdmiddotL
where d = pile diameter and L = length embedded below grade
If the magnitude of side friction is uniform and constant with depth the side capacity is simply
C-2
119928119956 = 119943119953 middot 119912119956 C1
Moreover however many piles extend through multiple layers and a summation of unit side
frictions acting on various pile segments must be tabulated over the length of the pile as
suggested by Figure C1
Figure C1 Components of Axial Pile Capacity
The unit-end bearing resistance (qb) acts over the base of the pile tip where the area of a circular pile is
given by Ab = πmiddotd24 For piles in compression loading the base capacity is determined from Equation
C2
119876119887 = 119902119887 middot 119860119887 C2
and for piles in tension (or uplift) it is normally taken that Qb = 0
C-3
C112 Pile Unit Side Friction
Several different approaches can be adopted for evaluating the unit pile side friction (fp) prior to full-
scale load testing and construction (Poulos amp Davis 1980 ONeill 2001) The most common types
including the alpha and beta methods as summarized in Tables 1 and 2 respectively In the alpha
method an empirical coefficient (α) is applied to the undrained shear strength (su) of clay soils to
obtain
119891119901 = 120572 middot 119904119906 C3
An illustrative example of alpha curves is shown in Figure C2 A difficulty with the alpha
method is the evolution of many variants and changes to the expressions for its estimation It is
also restricted in its use for specific types of piles in clays and fine-grained soils
C-4
Figure C2 Selected alpha relationships as function of the undrained shear strength
Table C1 Alpha Methods for Axial Pile Side Friction where fp = αmiddotsu
C-5
Reference
Source
Alpha Equation Remarks
Tomlinson
(1957)
2716801102
uu ssAll pile types (steel
concrete timber)
Note su in ksf
Tomlinson
(1957)
2616201102
uu ssConcrete piles
Note su in ksf
McClelland
(1974)
Kerisel α = 07 - 031middotln(su)
Peck α = 10 - 02middotsu
Woodward α = 091middot(su)091
Driven piles in clay
Note su in ksf
Semple
(1980) )1(511
)1(1
OCR
OCR Summary from 9 series of
pile load tests
American
Petroleum
Institute (API
1981)
For suσvo le 035 α = 10
For suσvo gt 080 α = 05
Otherwise α = 1 + 1111middot (035 - suσvo)
Driven steel pipe piles
Tomlinson
(1986)
1 α = 55 (su)-091
2 α = 785 (su)-102
3 α = 605 (su)-100
where 75 lt su (kPa) lt 210
1 Driven piles
2 Bored piles
3 Driven piles in till with L
lt 10middotd
API (1987) 1 α = 10 for su lt 25 kPa Driven piles in clay other
than Gulf of Mexico
C-6
2 α = 125 - 001middotsu for 25 lt su lt 75 kPa
3 α = 050 for su gt 75 kPa
Kulhawy amp
Jackson
(1989)
α = 021+026middot(σatm su) le 1 Analyses of 106 drilled
shaft foundations
Table C1 Continued
API (1989) 05 For suσ vo le 1 α =
su radic frasl prime σvo
minus025 su For suσ vo gt 1 α = 05 ∙ ( frasl prime ) σvo
Driven steel pipe piles
Kulhawy amp
Jackson
(1989)
OCR
OCR
sin50
tan)sin1( sin
Λ = (1-CsCc) asymp 080 for
insensitive clays
Nowacki et al
(1996)
minus05 su For suσvo le 07 α = 05 ∙ ( frasl prime ) σvo
minus02 su For suσvo gt 07 α = 055 ∙ ( frasl prime ) σvo
Driven pile foundations
Kolk and van
der Velde
(1996)
α =09middot FL middot (suσvo)-03 lt 10
FL = [ (L - z)d]-02 = length term
L = pile length
Note su obtained from
UU lab tests
Applicable to driven piles
in clays
C-7
Knappett amp α = 1 for su le 30 kPa Non-displacement piles in
Craig (2012) fine-grained soils α = 116 - (su185) for 30 kPa le su le 150 kPa
α = 035 for su gt 150 kPa
Knappett amp
Craig (2012) α = 055 ∙ 02 40
( ) (LfraslD)
∙ su ( frasl prime σvo
minus03
) Displacement piles in fine-
grained soils
z = depth at point considered
d = outside diameter of pile
Karlsrud et al
(2005 NGI
Method)
α = function (suσvo and plasticity index) Driven pilings
Fleming et al
(2009) 05 minus05 su su For su σvo le 1 120572 = ( frasl prime ) ∙ ( frasl prime )
σvo NC σvo
05 minus025 su su For su σvo gt 1 α = ( frasl prime ) ∙ ( frasl prime ) σvo NC σvo
For driven piles in clay
where (suσvo)NC is the
normalized shear strength
for normally-consolidated
clay
Brown et al
(2010)
α = 0 for 0 lt z le 5 feet
α = 055 for z gt 5 feet and (suσatm) le 15
α = 055-01middot(suσatm -15) for 15le (suσatm) le 25
Drilled Shafts and Bored
Piles
Table C1 Continued
C-8
OCRK )sin1(0
Karlsrud
(2012)
α = function (su σvo and plasticity index) Driven pilings
Note as reported by Karlsrud (2012)
In the beta method the coefficient β is applied to the effective overburden stress (σvo) at the point of
concern along the pile length
prime = 120573 C4 119891119901 middot 120590119907119900
Table C2 Beta Methods for Axial Pile Side Friction where fp = βmiddotσvo
Reference Source Beta Equation Remarks
Burland (1973) β = (1-sinϕ) middot tan ϕ Piles in NC clays
Meyerhof (1976) β = Kmiddottanδ where K = K0 = for NC clays
K = 15middotK0 for OC clays
Poulos amp Davis
(1980)
β = (1-sinϕ) middot tan ϕ middot OCR05 Piles in OC stiff clays
Table C2 Continued
C-9
Kulhawy amp Jackson
(1989)
β = (1-sinϕ) middot tan ϕ middot OCRsinϕ Piles in quartz sands and insensitive
clays
ONeill (2001) β = (1-sinϕ) middot tan θ middot OCRsinϕ Drilled shafts where θ = (δϕ)middotϕ and
δ = interface friction between pile and soil
Fleming et al
(2009)
β = Kmiddottanδ K = lateral stress coefficient and δ =
friction angle between soil and pile
material
Karlsrud (2012) β = fctn (OCR and PI) PI = plasticity index of the clay ()
Mayne and Niazi
(2017)
β = CM middot CK middot K0 middot tanϕ
where K0 = (1-sinϕ) middot OCRsinϕ for soils
that are virgin loaded then unloaded
CM = pile material factor = 1 (drilled
augered) 09 (prestressed or precast
concrete) 08 (timber) and 07
(steel) and CK = pile installation factor
= 09 (bored or augered) 10 (low
displacement eg H-pile or open-end
pipe) and 11 (driven high
displacement eg prestressed
concrete closed-end pipe)
As the beta approach applies to all types of soils (gravels sands silts and clays) and more or
less has remained unchanged since its advent circa 1970 it has been selected for further
discussion herein In the direct form for calculation of unit pile side friction the expression is
given by
119874119862119877119904119894119899120601 prime prime 119891119901 = 119862119872 ∙ 119862119870 ∙ (1 minus 119904119894119899120601prime) middot ∙ 119905119886119899120601prime ∙ 120590119907119900 C5
C-10
where CM is a soil-pile interface friction coefficient and CK = pile installation factor Values of CM are in
the range 07 lt CM lt 10 depending upon pile material (steel wood concrete) and ranges for CK vary
09 lt CK lt 11 depending upon method of installation (auger drill driven) as detailed in Table 2
The value of K0 is limited to the passive stress coefficient which for the simple Rankine case is given by
KP = (1+sin ϕ) (1 - sin ϕ ) Thus there is a maximum value of overconsolidation ratio for which
Equation C5 applies given by OCRlimit = [(1+sin ϕ ) (1 - sin ϕ )2] (1sinϕ)
The corresponding graph in Figure C3 shows the relationship for β in terms of ϕ and OCR
Figure C3 Pile side friction coefficient β in terms of effective friction angle and overconsolidation ratio
C113 Pile Unit End Bearing
C1131 Theoretical Considerations
The evaluation of the end bearing resistance of pile foundations is commonly determined using
limit plasticity theory where
prime Drained loading = C6 119902119887 119873119902 middot 120590119907119900
C-11
Undrained loading 119902119887 = 119873119888 middot 119904119906 C7
where the value of σvo is calculated at the depth z = L and the value of su is the average undrained shear
strength from z = L to a depth z = L + d beneath the pile tip For a circular pile the limit plasticity solution
for undrained loading gives (Vesic 1977) a value Nc = 933 while a deep strip foundation would employ a
value of Nc = 824 For drained loading the expression for Nq for a deep foundation is given by (Vesic
1977)
)]arctan()sin1(tan21[)](tan1[sin1
sin1)tanexp( 2 BLABNq
C8
where A and B are the pile plan dimensions (for a circular pile A = B) and L = pile length For a
circular pile this can be approximated by
asymp 077(120601prime75deg) 119873119902 C9
over a range of effective friction angles 20deg le ϕ le 45deg
C1132 Practical Considerations
For drained loading of sands the full calculated end bearing capacity will never be realized because it
would require the pile to move a distance equal to its diameter This is beyond practical use and
therefore the limit plasticity solutions must be clipped to a fraction of the calculated value Another
reason for using a reduced end-bearing capacity is due to strain incompatibility since the side capacity is
mobilized early while the end-bearing is engaged much later So to have compatible values of Qs and
Qb the qb must be reduced using the following guidelines (Randolph 2003 Mayne 2007)
prime prime C10 119902119887 = 119873119902 ∙ 120590119907119900 ∙ 119891119909
where fx = strain incompatibility factor
fx = 010 for drilled shafts augered cast-in-place and bored piles
fx = 020 for driven low displacement piles (opened-ended steel pipe and H-piles)
fx = 030 for driven high-displacement piles (ie solid piles PSC and closed-ended pipe)
C-12
For undrained loading of circular piles in compression no reduction of end-bearing is necessary
therefore
119902119887 = 933 middot 119904119906 C11
C12 Direct CPT Methods for Pile Capacity
In the direct CPT method the penetrometer readings are scaled directly via specified algorithms to
obtain the pile unit side friction and end-bearing as depicted in Figure C4 As many as 40 different
direct CPT methods have been developed over the past five decades as summarized by Niazi amp Mayne
(2013) Starting circa 1970 many of these early methods relied on hand-recorded data from field
mechanical CPTs where only qc data were obtained at 20 cm intervals or later with mechanical readings
of both qc and fs using special sets of inner and outer rods that recorded vertical load for tip and loads
for tip plus sleeve in alternating increments Also early pile load test data were often obtained from
top-down measurements of load-displacement
C-13
Undrained qb = Nc su
Qside = S (fp DAs)
QTotal = Qs + Qb - Wp
fp = cmck Ko svorsquo tanrsquo
Qbase = qb Ab
Drained qb = Nq svorsquo
Method Two Rational or ldquoIndirectrdquo Method
Method OneldquoDirectrdquo CPT Method
(Scaled Pile)
fp = fctn (soil type pile
type qt fs and u2)
qb = fctn (soil type qt-u2)
qb = unit end bearing
unit sidefriction fp
OCR su Ko t DR rsquo
AXIAL PILE CAPACITY FROM CPT READINGS
Figure C4 Concept of Direct versus Rational Method for CPT evaluation of axial pile capacity
Beginning in the mid-1990s as the modern electric piezocone (CPTu) was implemented the newer
equipment offered improved data and better resolution because of the additional reading of porewater
pressures and correction of raw measured qc to total resistance qt In addition the use of electronic
digital data collection and field computers proved superior in field measurements and recordings As a
consequence several reliable CPT methods for axial pile capacity have been developed Moreover
parallel improvements in full-scale pile load testing occurred and now include modern strain gage
instrumentation digital data recording and automated testing procedures Also the testing can be
single direction (compression or tension) or bi-directional as in the Osterberg cell
C-14
Table C3 provides a selection of recent direct CPT methods that have become available over the
past two decades A number of these (namely ICP NGI UWA and Fugro) were funded by the
offshore industry because of the growth of oil amp gas reserves and windfarm installations thus
necessitating increased concerns on risk probability and reliability in the site investigations for
offshore platforms and design of large driven monopile foundations These direct CPT methods
are often statistically based on large datasets compiled from full-scale load tests made
worldwide (eg Schneider et al 2008)
Table C3 Selection of Direct CPT Methods for Axial Pile Capacity
Method Pile Types Soil Types References CPT
data
Additional
parameters
needed
Unicone Method all types Sands silts
clays
Eslami and
Fellenius
(1997)
Fellenius
(2009)
qt fs u2
KTRI = Kajima
Technical
Research
Institute
all types Sands
mixed clays
Takesue et al
(1998)
fs and u2 No guidance given
on end bearing
resistance
Imperial College
Procedure (ICP)
OE and CE Sands Chow et al
(1997 PhD)
Jardine et al
(2005)
qt Interface friction
angle (δ) from
ring shear tests
Correlated to
mean grain size
(D50)
C-15
OE and CE Clays Jardine et al
(2005)
qt Interface friction
angle (δ)
correlated to PI
Table C3 Continued
NGI Method
(Norwegian
Geotechnical
Institute)
OE and CE Sands Clausen et al
(2005)
qt DR from qt
Clays a Alpha
method
(Karlsrud et al
2005 2012)
b Beta
method
(Karlsrud
2012)
qt for su
qt for
OCR
PI = plasticity
index ()
PI = plasticity
index ()
Fugro Method OE and CE Sands Kolk et al
(2005)
qt
Clays Van Dijk and
Kolk (2011)
qt
OE and CE Sands Lehane et al
(2005)
qt IFR = infilling ratio
related to amount
of plugging
C-16
Purdue LFRD Drilled
shafts
Sands Basu amp
Salgado
(2012)
qt LRFD = load
resistance
factored design
Enhanced Various Sands silts Niazi and qt u2 Based on 330 load
Unicone Method clays and Mayne (2015 and fs tests including
mixed soils 2016) bored augered
jacked and driven
piles
UWA (Univ of
Western
Ra = pile
roughness
Australia) Clays Lehane et al
(2012)
qt Interface friction
angle (δ) from
ring shear tests
and also method
without δ
HKU Method
(Hong Kong
Univ)
OE and CE
(end
resistance
only)
Sands Yu and Yang
(2012)
qt End bearing only
PLR = plug length
ratio
Note PLR can be
estimated from
OE inner diam
Table C3 Continued
Note previously called Marine Technical Directorate (MTD)
C-17
Many of the offshore CPT methods relate to driven piles either in sand or clay as that is
common deep foundation for those purposes Of particular interest are the Unicone and
Modified Unicone Methods since they use all three readings of the piezocone (qt fs and u2)
and address a variety of pile foundation types
C13 Modified UniCone Method
The modified UniCone Method was developed based on a total 330 pile load tests which were
associated with SCPTu data during their site investigations (Niazi and Mayne 2015 2016) This
represents a threefold increase over the original UniCone database that was built upon data
from 106 pile load tests (Eslami amp Fellenius 1997)
For the original UniCone algorithms use is made of the effective cone resistance (qE)
119902119864 = 119902119905 minus 1199062 C12
and a chart of qE vs fs provided an approximate soil classification in five distinct groups as shown by
Figure C5 Later in the modified approach a better delineation of the larger dataset gave soil
subclassifications as indicated by Figure C6
C-18
Figure C5 Soil behavior type using CPT via original UniCone
Figure C6 Soil behavior type using CPT via Modified UniCone
C-19
In the modified approach the 9-zone normalized soil behavioral type (SBTn) is ascertained using CPT
data in conjunction with the Robertson (2009) charts as shown in Figure C7 As discussed in Appendix
A and B the SBTn method uses the normalized cone resistance (Qtn) normalized sleeve friction (Fr())
and CPT material index Ic This permits a much wider range in the calculated pile side friction because fp
is related as a continuous curve with Ic rather than only 5 values that are assigned in the original
scheme
The pile unit side friction (fp) is obtained from qE and the CPT material index Ic using the following
expression at each elevation along the sides of the pile
10(0732 middot 119868119888 minus 3605) fp middot middot C13 = 119902119864 middot 120579119875119879 middot 120579119879119862 120579119877119860119879119864
C-20
Figure C7 Nine-zone soil behavioral soil type using normalized piezocone parameters
(after Robertson 2009 Mayne 2014)
where θPT = coefficient for pile type (θPT = 084 for bored piles 102 for jacked piles 113 for driven
piles) θTC = coefficient for loading direction (θTC = 111 for compression and 085 for tension) and θRATE
= rate coefficient applied to soils in SBT zones 1 through 7 (θRATE = 109 for constant rate of penetration
tests and 097 for maintained load tests) Since CPT provides data at regular intervals of 2 cm to 5 cm
along the sides of the pile the average fp from z = 0 to z = L can be used directly in Equation 1 to obtain
the shaft capacity Qs
C-21
The pile end bearing resistance is obtained from
middot 10(0325middot 119868119888minus1218) 119902119887 = 119902119864 C14
where qE is averaged in the vicinity of the pile tip Figure C8 shows the Modified Unicone Method in
graphical format For sensitive clays of zone 1 please see additional discussions by Niazi amp Mayne
(2016)
Figure C8 Summary of Modified UniCone Method for direct CPT assessment of axial pile
capacity
C-22
C14 Axial Pile Displacement
The movement of pile foundations can be assessed using elastic continuum theory (Poulos amp
Davis 1980 Randolph 2003 Mayne amp Niazi 2017) These relationships have been developed
using finite element analyses boundary elements and analytical closed-form solutions For the
latter approach the top displacement of a rigid pile subjected to a vertical force is shown in
Figure C9 This gives the movement at the top of a pile subjected to either compression or
tension (uplift) loading Also the percentage of axial load transferred to the pile toe is
determined For piles extending through various soil layers the elastic solution can be
implemented by using a set of stacked pile segments each with its own stiffness as
represented by a soil Youngs modulus
For pile groups the use of computer software is recommended Several available programs can handle
pile groups under axial and lateral moment loading such as DEFPIG (Univ Sydney) and PIGLET (Univ
Western Australia) A full listing of pile foundation software is given at the Geotechnical amp
GeoEnvironmental Service Directory wwwggsdcom
Ps
= Sh
aft
Load
sL
t
tEd
IPw
Load Transfer to Base
)]1)((5ln[
)(
)1(1
1
2 vdL
dLFI
E
ECT
211
IF
P
P CT
t
b
Base Load = Pb
Randolph Elastic Solution
Ground Surface
Top Load Pt = Ps + Pb
Rigid Pile Response
wt = displacementd = diameter
L = length
Es = Elastic Soil Modulus
Es0 (at z = 0) surface
EsM (at z = L2)mid-length
EsL (at z = L)full length
Load Transfer to Shaft
E = EsMEsL = Gibson parameterE = 1 for homogeneous case (constant E)E = 05 for pure Gibson case (Es0 = 0)
where FCT = load direction (= 1 compression 0 tension)
21
IF
P
P CT
t
b
z = depth
= Poissons ratio
Tension
Compression
C-23
Figure C9 Elastic continuum solution for axial pile response under compression and tension
loading
C-24
C15 Acknowledgements
The author appreciates the help of Fernando Illingworth of PonsEcuador Meena Viswanth of
GeoSyntecAtlanta and Dr Marco Uzielli of GeoRiskItaly in helping to collect and process the data
Funding support was provided by David Woeller of ConeTec Richmond BC
C-25
- Structure Bookmarks
-
- CONE PENETRATION TEST DESIGN GUIDE FOR STATE GEOTECHNICAL ENGINEERS
- TABLE OF CONTENTS
- LIST OF TABLES
- CHAPTER 1 INTRODUCTION
- CHAPTER 2 DIRECT CPT METHOD FOR SOIL CHARACTERIZATION
- CHAPTER 3 DIRECT CPT METHOD FOR SHALLOW FOUNDATIONS
- CHAPTER 4 DIRECT CPT METHOD FOR DEEP FOUNDATIONS
- REFERENCES
- APPENDIX A
- APPENDIX B
- APPENDIX C
-
CHAPTER 1 INTRODUCTION
The purpose of this manual is to provide an analysis of soil characterization shallow footings and deep
foundations using direct cone penetration testing (CPT) methods Geotechnical site characterization is
important for evaluating soil parameters that will be used in the analysis and design of foundations
retaining walls embankments and situations involving slope stability Common practice is to determine
these soil parameters through conventional lab and in-situ testing An alternative method uses CPT
readings of cone tip resistance (qt) sleeve friction ( fs) and pore pressure (u2) directly to determine these
parameters such as unit weight effective friction angle undrained shear strength and many others
(Figure 1) Direct CPT methods are provided for these parameters which will be used in the designs for
shallow and deep foundations
Figure 1 Geoparameters determined from CPT
Designing shallow foundations is typically done in a two-part process determining the bearing capacity
and expected settlement (commonly referred to as displacement) of the soil to approximate the
required size and shape of a foundation The older traditional methods are no longer required with
many approaches existing for using CPT directly in the design of shallow foundations Results from these
methods can provide a direct assessment of bearing capacity andor settlement A specific approach to
the direct method has been recommended and tested using a database of 166 full-scale field load tests
(Figure 2) A one-part process is used to scale the measured cone penetrometer readings (ie
measured cone tip resistance sleeve friction and pore water pressure) to obtain the bearing capacity
and settlement of the soil A step-by-step procedure has been created to transition from the CPT data to
bearing capacity with settlement accounted for The steps consist of estimating a design footing width
1
and length while using the process in the Soil Characterization section to determine soil parameters
directly from the CPT data
Figure 2 Conventional method for shallow foundation design compared to direct CPT method
There are upwards of 40 different direct CPT methods that have been developed over the past five
decades to determine the axial compression capacity of a piling foundation Earlier direct CPT methods
relied on hand-recorded information where mechanical-type CPT cone tip resistance data would be
collected at 20 cm intervals
Herein the method recommended for deep foundation design is the Modified UniCone method which
uses all three readings of the modern electronic piezocone penetrometer (CPTu) while addressing a
variety of pile foundation types (Figure 3) The modified UniCone Method is based on a total of 330 pile
load tests (three times the original UniCone database) that were associated with SCPTu data Two
computer software programs have been recommended for analyzing pile movements andor
settlements
2
Figure 3 Direct CPT evaluation of axial pile capacity
3
Symbol Parameter
γt Soil total unit weight
Ic CPT material index
SBT Soil behavior type (SBT)
ʹ σp Preconsolidation stress
YSR Yield stress ratio
ϕʹ Effective friction angle
Ko Lateral stress coefficient
su Undrained shear strength
Dʹ Constrained modulus
Eʹ Drained Youngrsquos modulus
CHAPTER 2 DIRECT CPT METHOD FOR SOIL
CHARACTERIZATION
21 INTRODUCTION
A direct CPT method for determining the value of each of various geoparameters shown in Table 1 is
provided The parameter Ic is used in the derivation of several sequential parameters This section of
the guide may be referred to as these parameters are used in calculations throughout the shallow
foundations and deep foundations sections
Table 1 Geoparameters calculated directly from CPT
4
Kʹ Bulk modulus
ks Subgrade reaction modulus
MR Resilient modulus
Gmax Small-strain shear modulus
k Coefficient of permeability
cv Coefficient of consolidation
Symbol Parameter Equation
ρt Mass Density ρt = γtga
where ga = 98 ms2
σvo Total Stress σvo asymp Σ (γti middot Δzi)
c Effective cohesion ʹ Empirical c asymp 003σp
In clays c asymp 01cu
Other minor parameters can be used in calculations of the geoparameters in Table 1 These minor parameters are provided in Table 2
Table 2 Minor Geoparameters
5
22 SOIL UNIT WEIGHT
The total soil unit weight can be estimated from CPT sleeve friction resistance as shown in Figure 4
(Mayne 2014) This method is not applicable to organic clays diatomaceous soils peats or sensitive
soils
Figure 4 Soil unit weight from CPT sleeve friction
Use Equation 1 to calculate the soils total unit weight
1 120574119905 = 120574119908 ∙ [122 + 015 ∙ ln (100 ∙ 119891119904 + 001)]
120590119886119905119898
Since soil unit weight is required for determining most geoparameters (including Soil Behavior Type)
estimating the soil type of the layers using the ldquorules of thumbrdquo method is a good first step before
determining more precise layering (Figure 5) Once a unit weight is determined for each noticeable layer
change these results can be used in later calculations such as ldquoCPT Material Indexrdquo To use the ldquorules
of thumbrdquo method some helpful guidelines are to assume sands are identified when qt gt 725 psi and u2
asymp uo while the presence of intact clays are prevalent when qt lt 725 psi and u2 gt uo The magnitude of
porewater pressures help to indicate intact clays such as soft (u2 asymp 2middotuo) firm (u2 asymp 4middotuo) stiff (u2 asymp 8middotuo)
and hard (u2 asymp 20middotuo) Fissured overconsolidated clays tend to have negative u2 values such that u2 lt 0
6
Figure 5 Approximate ldquorules of thumbrdquo method using CPT sounding from Wakota Bridge MN
23 CPT MATERIAL INDEX
The development of the CPT material index Ic has improved the initial classification of soil types and
calculation of soil parameters as shown in Table 1 To calculate Ic follow steps 1 and 2
231 Step 1 Normalized Sleeve Friction
Calculate Fr using Equation 2 if not provided with the CPT data gathered
2
7
119891119904 119865119903() = 100 ∙
(119902119905minus120590119907119900)
232 Step 2 Iteration
Iterate using Equations 3-5 to determine Ic by initially using n = 1 to calculate a starting value of Ic The
exponent n is soil-type dependent n = 1 (clays) n asymp 075 (silts) and n asymp 05 (sands) Iteration converges
quickly which is generally after the 3rd cycle
3 (119954119957minus120648119959119952)120648119938119957119950 = 119951 119928119957119951 prime (120648119959119952120648119938119957119950)
4
prime 120590119907119900 119899 = 0381 ∙ 119868119888 + 005 ( ) minus 015 120590119886119905119898
119899 le 10
5 119920119940 = radic[120785 120786120789 minus 119949119952119944119928119957119951]120784 + [120783 120784120784 + 119949119952119944119917119955]120784
24 SOIL BEHAVIOR TYPE (SBT)
Typically soil samples are not taken when using CPT The soil types are then inferred from the qt fs and
u2 readings To determine the types of soil from CPT data follow steps 1 and 2
241 Step 1
To determine the soil layers from CPT results calculate Ic by following the steps under section ldquoCPT
Material Indexrdquo After Ic has been determined through all specified depths use Figure 6 to classify the
type of soil by comparing each Ic value to normalized CPT readings (Fr and Qtn) from Equations 2 and 3
8
Figure 6 SBT zones using CPT Ic
242 Step 2
To determine if any of the soil layers contain ldquosensitive clays and siltsrdquo from zone 1 or ldquovery stiff
overconsolidated (OC) soilrdquo from zones 8 and 9 use Equations 6 and 7 If any soil layers are found within
zone 1 by Equation 6 then caution should be taken as these clays are prone to instability collapses and
difficulties in construction performance Very stiff OC sands to clayey sands of zone 8 (15 lt Fr lt 45)
and very stiff OC clays to silts of zone 9 (Fr gt 45) can be identified by Equation 7
6 Equation 6 errata This is an exponential expression (see Figure 5)
119876119905119899 lt 12119890119909119901(minus14 ∙ 119865119903)
7 119876119905119899 gt 0005(119865119903minus1)minus00003(119865119903minus1)2minus0002
1
Errata See terms and coefficients in Figure 5 above (numbers cannot be rounded off)
After each CPT reading has been assigned a zone from Figure 14 a visual representation can be made to
show the predominant layers by soil types (Figure A5)
25 EFFECTIVE STRESS FRICTION ANGLE
The effective friction angle (ϕ) is used to govern the strength for sands and clays where Equation 8 is
used for sands and Equation 9 is used for clays The value of ϕ for sands is derived from Ic so before ϕ
can be calculated refer to the iteration of Qtn under the ldquoCPT Material Indexrdquo section Once the values
9
Soil Type m
Fissured clays 11
Intact clays 10
Sensitive clays 09
Silt mixtures 085
Silty sands 080
Clean sands 072
Note m may be higher than 11 in fissured clays
of Ic and Qtn are calculated the type of soil can be determined from the section ldquoSoil Behavior Type
SBTrdquo The type of soil will dictate which equation to use for ϕ
Sands
8 120601prime(deg) = 176deg + 110deg log(119876119905119899)
Clays
9 119902119905minus120590119907119900 120601prime(deg) = 295deg ∙ 119861119902
0121 ∙ [0256 + 0336 ∙ 119861119902 + log ( )] prime 120590119907119900
Where = 119861119902 (1199062 minus 119906119900)(119902119905 minus 120590119907119900)
26 STRESS HISTORY
Determining the stress history can be characterized by an apparent yield stress ratio of the form
10 prime 120590119901
119884119878119877 = prime 120590119907119900
Where σp is defined as the preconsolidation stress or effective yield stress (Equation 10) The YSR is the
same equation as the more common overconsolidation ratio (OCR) but is now generalized to
accommodate mechanisms of preconsolidation such as ageing desiccation repeated cycles of wetting-
drying repeated freeze thaw cycles and other factors
11 prime prime 120590119901 = 120590119910 = 033(119902119905 minus 120590119907119900)119898prime
Where σp σy σvo and qt have units of kPa and the value of m depends on soil type with typical values shown in Table 3
Table 3 Soil type compared to exponent m
10
The value of m for non-fissured soils and inorganic clays and silts is derived from Ic (Figure 7) so before
m can be calculated (Equation 12) refer to the iteration of Ic under the ldquoCPT Material Indexrdquo section
Determine Ic for all soil layers before calculating m
12 028
119898prime = 1 minus 25 119868119888 1+( frasl ) 265
Once m is known for all soil layers the YSR can be determined
Figure 7 Yield stress exponent compared to CPT material index
A limiting value of YSR can be reached for clays and sands It can be calculated using Equation 13
13 (1 sin(emptyprime))
(1+sin(emptyprime)) 119884119878119877119897119894119898119894119905 = [ ]
(1minussin(emptyprime))2
27 LATERAL STRESS COEFFICIENT
The lateral stress coefficient Ko = σhoσvo commonly referred to as the at-rest condition is used to
represent the horizontal geostatic state of soil stress Ko can be calculated using Equation 14 but
11
14
119870119900 = (1 minus sin(emptyprime)) ∙ 119884119878119877sin( emptyprime)
sections such as ldquoStress Historyrdquo and ldquoEffective Stress Friction Anglerdquo will need to be referred to for
determining parameters ϕ and YSR
A maximum value for Ko can be determined by Equation 15
15 (1+sin(emptyprime))
119870119900119898119886119909 = = 1199051198861198992(45deg + emptyprime2) (1minussin(emptyprime))
28 UNDRAINED SHEAR STRENGTH
Loading on soils can result in fully drained partially drained or fully undrained conditions Sands
typically produce drained cases due to their high permeability but exceptions may occur in loose sands
during fast loading where the water does not have sufficient time to dissipate Clays exhibit low
permeability and thus often result in undrained loading cases when a load is applied quickly For soft-
firm clays the undrained shear strength (su) can be determined from CPT via Equation 16 where the
value of the bearing factor Nkt can be taken as 12
16 119902119905minus120590119907119900 119904119906 =
119873119896119905
In the case of remolded undrained shear strength from CPT su asymp fs
29 GROUND STIFFNESS AND SOIL MODULI
Determining the grounds stiffness can be measured from geoparameters such as the constrained
modulus (D) drained Youngrsquos modulus (E) bulk modulus (K) subgrade reaction modulus (ks) resilient
modulus (MR) and small-strain shear modulus (Gmax)
17 119863 prime asymp 5 ∙ (119902119905 minus 120590119907119900)
18
12
119864 prime = 119863 prime
11
19 119864prime
119870prime = [3∙(1minus2119907prime)]
The subgrade modulus (ks) is a combination of soil-structural properties which creates a parameter that
depends on the ground stiffness and the size of the loaded element
20 119864prime
119896119904 = [119889∙(1minus1199072)]
The resilient modulus MR applies to pavement analysis and design and can be calculated using Equation
21 where qt and fs are in MPa
21 053 119872119877 = (146119902119905 + 135511989111990414 + 236)244
The small strain shear modulus Gmax is a representation of the initial stiffness of all soils and rocks
Graphically it is the beginning portion of all stress-strain-strength curves for geomaterials Use Equation
22 to determine Gmax
22 119866119898119886119909 = 120588119905 ∙ 1198811199042
Where Vs = shear wave velocity (Equation 23) as measured by seismic cone penetration tests (SCPT) If
only cone penetration tests (CPT) or piezocone (CPTu) data are available the shear wave velocity may
be estimated from
23 03 119891119904 119881119904 (119898frasl119904) = [101 ∙ log(119902119905) minus 114]167 ∙ (100 ∙ )
119902119905
Where qt and fs have units of (kPa)
210 COEFFICIENT OF CONSOLIDATION
The coefficient of consolidation (cv) controls the rate that foundation and embankment settlements
occur By using results of CPT dissipation tests that measure the change in u2 readings over time the
value of cv can be determined Using Equation 24 cv can be determined based on piezocone dissipation
13
curves The equation below requires an estimate of the in-situ rigidity index (IR) of the soil If results of
SCPTU are available then IR may be determined from Gmax and qt per equation A38
24 0030∙(119886119888)2∙(119868119877)075
119888119907 = 11990550
Where ac = penetrometer radius (178 cm for 10-cm2 cone 220 cm for 15-cm2 cone) t50 = time to reach 50 dissipation IR = Gsu = undrained rigidity index G can be determined from the ldquoGround Stiffness and Soil Modulirdquo section
211 HYDRAULIC CONDUCTIVITY
The hydraulic conductivity also known as the coefficient of permeability (k) expresses the flow
characteristics of soils and has units of cms or feetday One method of calculation would be to use
Equation 25 where cv would need to be determined from ldquoCoefficient of Consolidationrdquo and D would
need to be determined from ldquoGround Stiffness and Soil Modulirdquo
25 119888119907∙120574119908 119896 =
119863prime
An alternative approach developed for soft normally-consolidated soils (Figure 8) is shown in Equation
26 where t50 (sec) values are used directly in assessing k in (cms)
26
14
125
119896 asymp ( 1
) 251∙11990550
Figure 8 k vs dissipation time for 50 consolidation (Mayne 2017)
212 EXAMPLE PROBLEMS
2121 Example 1 Direct CPT Methods for Geoparameters on Sands
Several geoparameters need to be determined based on the given CPT data collected for the South
Abutment of a bridge in Benton County MN (Figure 9) The groundwater table (GWT) was measured at
17 feet Determine all the geoparameters found in Table 4 at depths of 0 feet to 30 feet All sand layers
can be assumed ldquodrainedrdquo with ν = 02
15
Figure 9 CPT data from Benton County Minnesota for example problem 1
16
Table 4 Geoparameters evaluated for Case Example 1
Symbol Parameter Symbol Parameter
γt Soil total unit weight Ko Lateral stress coefficient
Ic CPT material index Dʹ Constrained modulus
SBT Soil behavior type (SBT) Eʹ Drained Youngrsquos modulus ʹ σp Preconsolidation stress Kʹ Bulk modulus
YSR Yield stress ratio MR Resilient modulus
ϕʹ Effective friction angle Gmax Small-strain shear modulus
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 10 Soil layers using ldquorules of thumbrdquo
γw = 6224 pcf
17
Layer 1
From Figure 10 fs = 17 psi taken as a representative value of the layer
17 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1204 pcf
145 psi
Layer 2
From Figure 10 fs = 17 psi taken as a representative value of the layer
17psi
γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1204 pcf 145 psi
Layer 3
From Figure 10 fs = 12 psi taken as a representative value of the layer
12 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 4
From Figure 10 fs = 7 psi taken as a representative value of the layer
CPT Material Index
7 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1121 pcf
145 psi
Layer 1
From Figure 10 fs = 17 psi and qc = 3500 psi
18
qt = qc + u2 ∙ (1 minus a) = 3500 psi + 3 psi ∙ (1 minus 08) = 3501 psi
= γt ∙ 6 feet = 1204 pcf ∙ 6 feet = 7225 psf = 50 psi σvo
fs 17 psi Fr() = 100 ∙ = 100 ∙ = 049
(qt minus σvo) (3501 psi minus 50 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 5 psi minus 0 psi = 50 psi
(qt minus σvo)σatm (3501 psi minus 50 psi )145 psi = = Qtn prime )n (σvoσatm (50 psi145 psi)n
prime σvo 50 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(049)]2
Qtn = 35329 n = 036 lt 10 Ic = 13
Layer 2
From Figure 10 fs = 17 psi and qc = 3500 psi
19
qt = qc + u2 ∙ (1 minus a) = 3500 psi + 0 psi ∙ (1 minus 08) = 3500 psi
= 7225 psf + γt ∙ 6 feet = 7225 psf + 1204 pcf ∙ 6 feet = 1445 psf = 100 psi σvo
fs 17 psi Fr() = 100 ∙ = 100 ∙ = 049
(qt minus σvo) (3500 psi minus 100 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 100 psi minus 0 psi = 100 psi
(qt minus σvo)σatm (3500 psi minus 100 psi )145 psi = = Qtn prime )n (σvoσatm (100 psi145 psi)n
prime σvo 100 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
47 minus log(Qtn)]2 + [122 + log(049)]2
Ic = radic[3
Qtn = 27947 n = 041 lt 10 Ic = 14
Layer 3
From Figure 10 fs = 12 psi and qc = 1500 psi
20
qt = qc + u2 ∙ (1 minus a) = 1500 psi + 0 psi ∙ (1 minus 08) = 1500 psi
= 1445 psf + γt ∙ 11 feet = 1445 psf + 1172 pcf ∙ 11 feet = 2734 psf = 190 psi σvo
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 081
(qt minus σvo) (1500 psi minus 190 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = γwater ∙ (z minus 119911119908) = 6224 pcf ∙ (23 feet minus 17 feet) = 3734 psf = 99 psi
prime σvo = σvo minus uo = 190 psi minus 99 psi = 164 psi
(qt minus σvo)σatm (1500 psi minus 190 psi )145 psi = = Qtn prime )n (σvoσatm (164 psi145 psi)n
prime σvo 164 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(081)]2
Qtn = 947 n = 06 lt 10 Ic = 19
21
Layer 4
From Figure 10 fs = 7 psi and qc = 1200 psi
qt = qc + u2 ∙ (1 minus a) = 1200 psi + 0 psi ∙ (1 minus 08) = 1200 psi
= 2734 psf + γt ∙ 11 feet = 2734 psf + 1121 pcf ∙ 7 feet = 3519 psf = 244 psi σvo
fs 7 psi Fr() = 100 ∙ = 100 ∙ = 060
(qt minus σvo) (1200 psi minus 244 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = γwater ∙ (z minus 119911119908) = 6224 pcf ∙ (30 feet minus 17 feet) = 8091 psf = 130 psi
prime σvo = σvo minus uo = 244 psi minus 130 psi = 188 psi
(qt minus σvo)σatm (1200 psi minus 244 psi )145 psi = = Qtn prime )n (σvoσatm (188 psi145 psi)n
prime σvo 188 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(060)]2
22
Qtn = 686 n = 064 lt 10 Ic = 19
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic Qtn and Fr the first layer is defined as a ldquoDrained Gravelly Sandrdquo from
Figure 11
Figure 11 Soil layer 1 using SBT method
23
Layer 2
Based on values of Ic Qtn and Fr the second layer is defined as a ldquoDrained Sandrdquo from Figure
12
Figure 12 Soil layer 2 using SBT method
24
Layer 3
Based on values of Ic Qtn and Fr the third layer is defined as a ldquoDrained Sandrdquo from Figure 13
Figure 13 Soil layer 3 using SBT method
25
Layer 4
Based on values of Ic Qtn and Fr the fourth layer is defined as a ldquoDrained Sandrdquo from Figure 14
Figure 14 Soil layer 4 using SBT method
Effective Stress Friction Angle
Layer 1
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(3533) = 456deg
Layer 2
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(2795) = 445deg
26
Layer 3
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(947) = 393deg
Layer 4
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(686) = 378deg
Stress History
Layer 1
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (13frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(24139 kPa minus 345 kPa)072 = 4717 kPa
= 684 psi
prime σp 684 psi YSR = = = 136
primeσvo 50 psi
Layer 2
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + (14 1 + ( frasl ) frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(34132 kPa minus 689 kPa)072 = 605 kPa
= 877 psi
prime σp 877 psi YSR = = = 87
primeσvo 100 psi
27
Layer 3
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + (19 1 + ( frasl ) frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(10342 kPa minus 131 kPa)072 = 254 kPa
= 368 psi
prime σp 368 psi YSR = = = 22
primeσvo 164 psi
Layer 4
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (19frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(8274 kPa minus 168 kPa)072 = 215 kPa = 312 psi
prime σp 312 psi YSR = = = 17
primeσvo 188 psi
Lateral Stress Coefficient
Layer 1
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(456deg)) ∙ 136sin( 456deg) = 18
Layer 2
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(445deg)) ∙ 87sin( 445deg) = 14
28
Dprime 17480 psi
Eprime = = = 15890 psi 11 11
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(393deg)) ∙ 22sin( 393deg) = 06
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(378deg)) ∙ 17sin( 378deg) = 05
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3501 psi minus 50 psi) = 17480 psi
Eprime 15890 psi
Kprime = = = 8828 psi [3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Layer 3
Layer 4
Ground Stiffness and Soil Moduli
Layer 1
Values of qt and fs need to be in MPa
MR = (146(241 MPa)053 + 1355(012 MPa)14 + 236)244 = 3416 MPa = 49575 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1172 kPa m Vs (mfrasls) = [101 ∙ log(24139 kPa) minus 114]167 ∙ (100 ∙ ) = 2745
24139 kPa s
Vs = 9012 fts
29
γt 1204 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 37 ∙ (9012 ) = 304 ∙ 106psf = 21000 psi s
Layer 2
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3500 psi minus 100 psi) = 17480 psi
Dprime 17480 psi Eprime = = = 15890 psi
11 11
Eprime 15890 psi Kprime = = = 8828 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(241 MPa)053 + 1355(012 MPa)14 + 236)244 = 3415 MPa = 49562 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1172 kPa m Vs (mfrasls) = [101 ∙ log(24133 kPa) minus 114]167 ∙ (100 ∙ ) = 2745
24133 kPa s
Vs = 9012 fts
γt 1204 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
30
2 ft Gmax = ρt ∙ Vs
2 = 37 ∙ (9012 ) = 304 ∙ 106psf = 21000psi s
Layer 3
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (1500 psi minus 190 psi) = 7405 psi
Dprime 7405 psi Eprime = = = 6732 psi
11 11
Eprime 6732 psi Kprime = = = 3740 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(103 MPa)053 + 1355(008 MPa)14 + 236)244 = 1506 MPa = 21854 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 827 kPa m Vs (mfrasls) = [101 ∙ log(10343 kPa) minus 114]167 ∙ (100 ∙ ) = 2611
10343 kPa s
Vs = 8566 fts
γt 1172 pcf slug ρt = = = 36
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 36 ∙ (8566 ) = 27 ∙ 106psf = 19000 psi s
Layer 4
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (1200 psi minus 244 psi) = 5878 psi
31
Dprime 5878 psi Eprime = = = 5343 psi
11 11
Eprime 5343 psi Kprime = = = 2969 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(83 MPa)053 + 1355(005 MPa)14 + 236)244 = 165 MPa = 16901 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 483 kPa m Vs (mfrasls) = [101 ∙ log(8274 kPa) minus 114]167 ∙ (100 ∙ ) = 2243
8274 kPa s
Vs = 7360 fts
γt 1121 pcf slug ρt = = = 35
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 35 ∙ (7360 ) = 389 ∙ 106psf = 13000 psi s
32
2122 Example 2 Direct CPT Methods for Geoparameters on Clay
Several geoparameters need to be determined based on the given CPT data collected for the South
Abutment (Figure 15) The groundwater table (GWT) was measured at 60 feet Determine all the
geoparameters found in Table 5 at depths of 0 feet to 42 feet All sand layers can be assumed ldquodrainedrdquo
ν = 02 All clay layers can assume ν = 049
33
Figure 15 CPT data from Minnesota for example problem 2
34
Table 5 Geoparameters
Symbol Parameter Symbol Parameter
γt Soil total unit weight Eʹ Drained Youngrsquos modulus
Ic CPT material index Kʹ Bulk modulus
SBT Soil behavior type (SBT) MR Resilient modulus
ʹ σp Preconsolidation stress Gmax Small-strain shear modulus
YSR Yield stress ratio su Undrained shear strength
ϕʹ Effective friction angle cv Coefficient of consolidation
Ko Lateral stress coefficient k Hydraulic conductivity
Dʹ Constrained modulus
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
35
Figure 16 Soil layers using ldquorules of thumbrdquo
Unit weight of water γw = 6224 pcf
Layer 1
From Figure 16 fs = 13 psi taken as a representative value of the layer
13 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1179 pcf
145 psi
Layer 2
From Figure 16 fs = 12 psi taken as a representative value of the layer
12 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 3
From Figure 16 fs = 20 psi taken as a representative value of the layer
36
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
Layer 4
From Figure 16 fs = 2 psi taken as a representative value of the layer
2 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1004 pcf
145 psi
CPT Material Index
Layer 1
From Figure 10 fs = 13 psi and qc = 3000 psi
qt = qc + u2 ∙ (1 minus a) = 3000 psi + 0 psi ∙ (1 minus 08) = 3000 psi
σvo = γt ∙ 2 feet = 1179 pcf ∙ 2 feet = 235 psf = 16 psi
fs 13 psi Fr() = 100 ∙ = 100 ∙ = 043
(qt minus σvo) (3000 psi minus 16 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 16 psi minus 0 psi = 16 psi
(qt minus σvo)σatm (3000 psi minus 16 psi )145 psi = = Qtn prime )n (16 psi145 psi)n (σvoσatm
prime σvo 16 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(043)]2
37
Qtn = 4126 n = 032 lt 10 Ic = 121 therefore sand
Layer 2
From Figure 10 fs = 12 psi and qc = 250 psi
qt = qc + u2 ∙ (1 minus a) = 250 psi + 10 psi ∙ (1 minus 08) = 252 psi
σvo = 235 psf + γt ∙ 30 feet = 235 psf + 1172 pcf ∙ 30 feet = 3751 psf = 260 psi
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 53
(q minus σ ) (252 psi minus 260 psi) t vo
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 260 psi minus 0 psi = 260 psi
(qt minus σvo)σatm (250 psi minus 260 psi )145 psi = = Qtn prime (σvoσatm)n (260 psi145 psi)n
38
prime σvo 260 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(53)]2
Qtn = 87 n = 10 le 10 Ic = 32 therefore clay
Layer 3
From Figure 10 fs = 20 psi and qc = 4000 psi
qt = qc + u2 ∙ (1 minus a) = 4000 psi + 8 psi ∙ (1 minus 08) = 40016 psi
σvo = 3751 psf + γt ∙ 2 feet = 3751 psf + 1219 pcf ∙ 2 feet = 3995 psf = 277 psi
fs 16 psi Fr() = 100 ∙ = 100 ∙ = 054
(qt minus σvo) (30016 psi minus 277 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 277 psi minus 0 psi = 277 psi
(qt minus σvo)σatm (30016 minus 277)145 psi = = Qtn prime (σvoσatm)n (277 psi145 psi)n
39
primeσvo 277 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(054)]2
Qtn = 1962 n = 052 le 10 Ic = 15 therefore sand
Layer 4
From Figure 10 fs = 2 psi and qc = 250 psi
qt = qc + u2 ∙ (1 minus a) = 250 psi + 0 psi ∙ (1 minus 08) = 250 psi
= 3994 psf + γt ∙ 8 feet = 3994 psf + 1004 pcf ∙ 8 feet = 4798 psf = 333 psi σvo
fs 2 psi Fr() = 100 ∙ = 100 ∙ = 092
(qt minus σvo) (250 psi minus 333 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
uo = 0 psi
prime σvo = σvo minus uo = 333 psi minus 0 psi = 333 psi
(qt minus σvo)σatm (250 psi minus 333 psi )145 psi = = Qtn prime (σvoσatm)n (333 psi145 psi)n
40
prime σvo 333 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(092)]2
Qtn = 65 n = 10 lt 10 Ic = 29 therefore clayey silt
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic Qtn and Fr the first layer is defined as a ldquoDrained Sandrdquo from Figure 17
41
Figure 17 Soil layer 1 using SBT method
Layer 2
Based on values of Ic Qtn and Fr the second layer is defined as a ldquoUndrained Clayrdquo from Figure
18
42
Figure 18 Soil layer 2 using SBT method
43
Layer 3
Based on values of Ic Qtn and Fr the third layer is defined as a ldquoDrained Sandrdquo from Figure 19
Figure 19 Soil layer 3 using SBT method
44
Layer 4
Based on values of Ic Qtn and Fr the fourth layer is defined as a ldquoUndrained Silty Mixrdquo from
Figure 20
Figure 20 Soil layer 4 using SBT method
Effective Stress Friction Angle
Layer 1 (sand)
45
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(4126) = 464deg
Layer 2 (clay)
qt minus σvo
ϕprime(deg) = 295deg ∙ Bq0121 ∙ [0256 + 0336 ∙ Bq + log ( )]
primeσvo
Where
(u2 minus uo) (10 psi minus 0 psi) Bq = = = 004
(qt minus σvo) (252 psi minus 260 psi)
252 psi minus 260 psi ϕprime(deg) = 295deg ∙ 00440121 ∙ [0256 + 0336 ∙ 0044 + log ( )]
260 psi
ϕprime(deg) = 245deg
Layer 3 (sand)
ϕprime(deg) = 176deg + 110deg log(Qtn) = 176deg + 110deg log(1962) = 428deg
Layer 4 (clay)
qt minus σvo
ϕprime(deg) = 295deg ∙ Bq0121 ∙ [0256 + 0336 ∙ Bq + log ( )]
primeσvo
Where
(u2 minus uo) (5 psi minus 0 psi) Bq = = = 002
(qt minus σvo) (251 psi minus 333 psi)
252 psi minus 333 psi ϕprime(deg) = 295deg ∙ 0020121 ∙ [0256 + 0336 ∙ 002 + log ( )]
333 psi
ϕprime(deg) = 265deg
Stress History
46
Layer 1
028 028
prime m = 1 minus = 1 minus = 072 25 25 1 + (
Icfrasl ) 1 + (12frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(3000 psi minus 33)072 = 1051 psi
prime σp 1051 psi YSR = = = 642
primeσvo 16 psi
Layer 2
028 028
prime m = 1 minus = 1 minus = 099 25 25 1 + (
Icfrasl ) 1 + (32frasl ) 265 265
prime prime σp = σy = 033(qt minus σvo)mprime = 033(252 psi minus 260)099 = 734 psi
prime σp 734 psi YSR = = = 28
primeσvo 260 psi
Layer 3
028 028
prime m = 1 minus = 1 minus = 072 25 25 Ic 1 + ( frasl ) 1 + (15frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(4002 psi minus 277)072 = 1288 psi
47
prime σp 1288 psi YSR = = = 46
primeσvo 277 psi
Layer 4
028 028
prime m = 1 minus = 1 minus = 097 25 25 Ic 1 + ( frasl ) 1 + (29frasl ) 265 265
)mprime prime prime σp = σy = 033(qt minus σvo = 033(250 psi minus 333)097 = 626 psi
prime σp 626 psi YSR = = = 19
primeσvo 333 psi
Lateral Stress Coefficient
Layer 1
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(464deg)) ∙ 642sin( 464deg) = 56
Layer 2
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(245deg)) ∙ 28sin( 245deg) = 090
Layer 3
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(428deg)) ∙ 46sin( 428deg) = 091
Layer 4
Ko = (1 minus sin(emptyprime)) ∙ YSRsin( emptyprime) = (1 minus sin(266deg)) ∙ 19sin( 266deg) = 073
48
Ground Stiffness and Soil Moduli
Layer 1
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (3000 psi minus 16 psi) = 14991 psi
Dprime 14991 psi Eprime = = = 13629 psi
11 11
Eprime 13629 psi Kprime = = = 7571 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(207 MPa)053 + 1355(009 MPa)14 + 236)244 = 2816 MPa = 40873 psi
fs
03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 896 kPa m Vs (mfrasls) = [101 ∙ log(20685 kPa) minus 114]167 ∙ (100 ∙ ) = 2564
20685 kPa s
Vs = 8412 fts
γt 1179 pcf slug ρt = = = 37
ga 322 ftfrasls2 ft3
2 ft 2 Gmax = ρt ∙ Vs = 37 ∙ (8412 ) = 260 ∙ 106psf = 17992 psi
s
Layer 2
49
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (252 psi minus 260 psi) = 11298 psi
Dprime 11298 psi Eprime = = = 10271 psi
11 11
Eprime 10271 psi Kprime = = = 17118 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(049))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(17 MPa)053 + 1355(008 MPa)14 + 236)244 = 443 MPa = 64309 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 827 kPa m Vs (mfrasls) = [101 ∙ log(17375 kPa) minus 114]167 ∙ (100 ∙ ) = 2646
17375 kPa s
Vs = 8680 fts
γt 1172 pcf slug ρt = = = 36
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 36 ∙ (8680 ) = 274 ∙ 106psf = 19036 psi s
Layer 3
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (4002 psi minus 277 psi) = 19869 psi
Dprime 19869 psi Eprime = = = 18063 psi
11 11
50
Eprime 18063 psi Kprime = = = 10035 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(02))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
MR = (146(276 MPa)053 + 1355(014 MPa)14 + 236)244 = 4017 MPa = 58309 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 1379 kPa m Vs (mfrasls) = [101 ∙ log(27591 kPa) minus 1379]167 ∙ (100 ∙ ) = 2854
27591 kPa s
Vs = 9363 fts
γt 121 pcf slug ρt = = = 38
ga 322 ftfrasls2 ft3
2 ft Gmax = ρt ∙ Vs
2 = 38 ∙ (9363 ) = 33 ∙ 106psf = 23050 psi s
Layer 4
Dprime asymp 5 ∙ (qt minus σvo) = 5 ∙ (250 psi minus 333 psi) = 1088 psi
Dprime 1088 psi Eprime = = = 989 psi
11 11
Eprime 989 psi Kprime = = = 16490 psi
[3 ∙ (1 minus 2vprime)] [3 ∙ (1 minus 2(049))]
053 MR = (146qt + 1355fs14 + 236)244
Values of qt and fs need to be in MPa
51
MR = (146(17 MPa)053 + 1355(001 MPa)14 + 236)244 = 360 MPa = 52189 psi
fs 03
Vs (mfrasls) = [101 ∙ log(qt) minus 114]167 ∙ (100 ∙ ) qt
Values of qt and fs need to be in kPa
03 138 kPa m Vs (mfrasls) = [101 ∙ log(17238 kPa) minus 114]167 ∙ (100 ∙ ) = 1545
17238 kPa s
Vs = 5069 fts
γt 1004 pcf slug ρt = = = 31
ga 322 ftfrasls2 ft3
2 ft 2 Gmax = ρt ∙ Vs = 31 ∙ (5069 ) = 80 ∙ 105psf = 5565 psi
s
Undrained Shear Strength
Layer 2
qt minus σv0 250 psi minus 26 psi su = = = 187 psi
12 12
Layer 4
qt minus σv0 250 psi minus 333 psi su = = = 181 psi
12 12
Coefficient of Consolidation
52
Example dissipation data are shown in Figure 21 Example calculations are provided
Figure 21 Dissipation t50 data
Layer 1
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (7197)075
cv = = = 359 cmsec 1 sec t50
Layer 2
53
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (94392)075
cv = = = 0008 cmsec 3000 sec t50
Layer 3
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (8303)075
cv = = = 665 cmsec 06 sec t50
Layer 4
)2 ∙ (IR)075 0030 ∙ (ac 0030 ∙ (220 cm)2 ∙ (26714)075
cv = = = 087 cmsec 11 sec t50
Hydraulic Conductivity
Layer 1
cm 1 psi 359 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 10 ∙ 10minus4 cmsec Dprime 14984 psi
Layer 2
cm 1 psi 0008 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 73 ∙ 10minus7 cmsec Dprime 11379 psi
Layer 3
cm 1 psi 665 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 20 ∙ 10minus5 cmsec Dprime 148650 psi
54
Layer 4
cm 1 psi 087 ∙ 6224 pcf ∙ cv ∙ γw sec 144 pcf
k = = = 18 ∙ 10minus4 cmsec Dprime 10861 psi
55
CHAPTER 3 DIRECT CPT METHOD FOR SHALLOW
FOUNDATIONS
31 PROCEDURE
Shallow foundation analysis is typically done in a two-part traditional procedure The traditional
techniques are no longer required as a direct CPT method for square rectangular and circular shallow
footings is available (Figure 22) This process has the soil types grouped into four main categories sands
silts fissured clays and intact clays When determining soil types for each design it is believed footings
on sands and silts act in a fully drained manner while intact clays act in an undrained manner under
conditions of constant volume In order to determine the vertical stress-displacement-capacity of
square rectangular and circular shallow footings follow the steps provided towards the solution given
by Equation 27
Figure 22 Direct CPT method for shallow foundations
Equation 27 may be used to calculate all footing stresses from zero to the bearing capacity (qmax) To
calculate qmax for a sized footing of width (B) length (L) and thickness (t) follow steps 1 through 7
provided The settlement (s) can be determined after the calculation of qmax by simply rearranging
56
Equation 27 where the allowable stress (qallow) is defined as qmax divided by the factor of safety (FS) For
shallow footings a FS value of 3 is commonly used in geotechnical engineering
120782120787 minus120782120785120786120787 119956 119923 119954119950119938119961 = 119945119956 ∙ 119954119957119951119942119957 ∙ (119913
) ∙ (119913
) 27 119950119938119961
2 119902119898119886119909 0345 1 frasl119865119878 119871 119904 = 119861 ∙ [ ∙ ∙ ( ) ] 28
119861 ℎ119904 119902119905119899119890119905
Where hs = the foundation soil formation parameter qtnet = the net corrected cone tip resistance
311 Step 1 Estimating Footing Dimensions
In Minnesota frost heave can have devastating effect on a shallow foundations It is common practice to
place a foundation bearing elevation below the expected maximum frost depth (roughly 45 to 6 feet
below ground level) With this assumption estimate a footing size (B x L) for design to obtain
representative data from CPT roughly 15middotB below the foundation depth (Df) as shown in Figure 23 The
CPT data collected will consist of
cone tip resistance (qc)
measured porewater pressure acting behind the cone tip (u2) as shown in Figure 24
sleeve friction (fs)
57
Figure 23 Direct CPT method introduction
Figure 24 Differentiation of porewater pressure measurement locations (Lunne et al 1997)
312 Step 2 Soil Characterization
The soil behavior type (SBT) and CPT material index (Ic) govern the value of the formation factor hs The
first step is following the steps in Soil Unit Weight to determine soil layering and γt values for each
layer To determine a representative unit weight (γsoil) for all the layers in the range of Df to 15∙B below
58
Df the user will need to use their engineering judgement on the definition of ldquorepresentative unit
weightrdquo This single value of unit weight is used in further calculations such as the total vertical soil stress (σvo) from Equation 29 and effective vertical stress (σvo) from Equation 30 Values of σvo and σvo
will need to be calculated at 15middotB below Df
120648119959119952 = sum(120632119956119952119946119949 ∙ 119963) 29
prime 120648119959119952 = 120648119959119952 minus 119958119952 30
Where z = Df +15middotB uo=γwatermiddot(z-zw)
The qc will need to be corrected using Equation 31 in the case of fine-grained soils that develop excess porewater pressure during cone penetration These values will be used in the following calculations
119954119957 = 119954119940 + 119958120784 ∙ (120783 minus 119938) 31
Where 119860119899 a = cone area ratio = eg MnDOT commonly uses a = 08 119860119888
119860119899 = cross-sectional area of load cell or shaft 119860119888 = projected area of the cone
The cone area ratio is determined based on the type of piezocone tip used during in-situ field testing
Manufacturer specifications should provide the measured net area ratio (a) for the particular cone
penetrometer as determined by calibration in a pressurized triaxial chamber
313 Step 2a Foundation soil formation parameter
With the representative value γsoil continue to follow the steps in CPT Material Index through Soil
Behavior Type (SBT) to better define the type of soil at depth 15∙B below Df Use the value of Ic
calculated at depth 15∙B below Df to calculate hs with Equation 36 This parameter is based on the soil
type with typical values shown in Figure 25 The data on silts and sands are considered fully drained
whereas the fissured clay subset may be partially drained to undrained
59
23 ℎ119904 = 28 minus
119868119888 15 32
1+( ) 24
Figure 25 Foundation soil formation parameter hs versus CPT material index Ic (Mayne 2017)
314 Step 3 Soil elastic modulus and Poissonrsquos ratio
Details about the soil such as its elastic modulus (Es) and Poissonrsquos ratio (ν) will be needed for further
calculations A representative value for the Es can be determined from in-situ field tests Values of ν can
be assigned as 02 for drained sands and as 05 for undrained loading cases involving clays (Jardine et al
1985 Burland 1989)
315 Step 4 Net cone tip resistance
Calculate qtnet the mean value of net cone tip resistance 15middotB below the foundation bearing elevation
using Equation 33
33 119902119905119899119890119905 = 119902119905 minus 120590119907119900
60
316 Step 5 Bearing capacity of the soil
Use the assumed B and L calculated hs and qtnet to determine the soils bearing capacity (qmax) from
Equation 34 Use Table 6 to calculate qmax by using the maximum allowable settlement ratio (sB)max
correlating to the soil type If hs is in between the given values interpolate to acquire (sB)max
Table 6 Bearing capacity defined by soil type
Type of Soil hs (sB)max
Clean Sands 058 12
Silts 112 10
Fissured Clays 147 7
Intact Clays 270 4
05 minus0345 119904 119871 119902119898119886119909 = ℎ119904 ∙ 119902119905119899119890119905 ∙ (119861
) ∙ (119861
) 34 119898119886119909
317 Step 6 Settlement
Settlement can be calculated directly using the results from Equation 35 A FS equal to 3 is common in
foundation engineering
2 119902119898119886119909 0345 1 frasl119865119878 119871 119904 = 119861 ∙ [ ∙ ∙ ( ) ] 35
119861 ℎ119904 119902119905119899119890119905
318 Step 7 Final Check
Check that the applied stress (q) is less than qmax using Equation 36 Repeat the process again with a
new B andor new L if q gt qmax
05 119904 ) 119902 = ℎ119904 ∙ 119902119905119899119890119905 ∙ (119861
minus0345 119871 ∙ ( )
11986136
61
32 EXAMPLE PROBLEMS
321 Example 3 Direct CPT Method on Sands
A footing size needs to be determined based on the given CPT data collected for the South Abutment
(Figure 27) The footing stress (q) was determined to be 8000 psf Estimate a footing size (B x L) and
determine the bearing capacity of the foundation (qmax) using the direct CPT method provided Also
determine the expected settlement based on the calculated bearing capacity
Figure 26 Diagram of footing profiles for Example 3
The soil elastic modulus determined from seismic CPT (SCPT) are shown in Table 7
Table 7 SCPT Results
Depth (feet) Es (tsf)
Bottom of layer
3 557
6 433
9 557
12 695
17 590
22 501
27 571
32 505
62
Figure 27 CPT data from Northern Minnesota
Solution
Step 1 Assume the footing is placed below the frost depth of 6 feet
Estimate L L = 50 feet = 600 inches
63
Estimate B B = 12 feet = 144 inches
Estimate footing thickness t t = 2 feet
Df = 6 feet = 72 inches
Df + 15middotB = 24 feet = 288 inches
Step 2 Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 28 Schematic of foundation design associated with CPT data
Layer 1
From Figure 28 fs = 20 psi taken as a representative value of the layer
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
64
Layer 2
From Figure 28 fs = 20 psi taken as a representative value of the layer
20psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
Layer 3
From Figure 28 fs = 8 psi taken as a representative value of the layer
8 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1134 pcf
145 psi
Using the unit weights calculated between Df and 15B determine a representative unit weight of the soil to calculate the total and effective soil stresses Based on layer 3 being the weakest supporting layer γsoil = 113 pcf
σvo = γsoil ∙ (119863119891 + 15 ∙ B) = 113 lbsfraslft3 ∙ 24 feet = 2721 psf = 189 psi
prime σvo = σvo minus uo = 2721 psf minus 0 psf = 2721 psf = 189 psi
Calculate the cone tip resistance between Df and Df + 15B below the foundation depth
qt = qc + u2 ∙ (1 minus a) = 1250 psi + 0 psi ∙ (1 minus 08) = 1000 psi
Step 2a CPT Material Index
Using Figure 28 the sleeve friction at 15B below the foundation depth is about 16 psi
65
fs 16 psi Fr() = 100 ∙ = 100 ∙ = 13
(qt minus σvo) (1250 psi minus 189 psi)
Iterate to solve for Ic Steps are not shown for brevity
(qt minus σvo)σatm (1250 psi minus 189 psi )145 psi = = Qtn prime (σvoσatm)n (189 psi145 psi)n
prime σvo 189 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Qtn = 703 n = 072 lt 10 Ic = 210
66
Figure 29 Soil type for example problem 3
Use the Ic value to determine the foundation soil formation parameter
Calculate the foundation soil formation parameter The tip resistance is about 1250 psi at 24 feet and
the cone area ratio was determined to be 08 from information provided by the manufacturer
23 23 hs = 28 minus = 28 minus = 078 = SandSilt 15 15 Ic 210
1 + ( ) 1 + ( ) 24 24
67
Step 3 Determine representative values of soil elastic modulus from soil testing and Poissons ratio The
soil elastic modulus was taken as the average value between depths of Df and 15middotB (6 feet and 24 feet)
433 + 557 + 695 + 590 + 501 Es = = 555 tsf = 708 psi
5
ldquoDrained sandsiltrdquo gives a ν = 020
Step 4 Calculate the net cone tip resistance
qtnet = qt minus σvo = 1250 psi minus 189 psi = 12311 psi
Step 5 Calculate the bearing capacity of the sand
In drained sandssilts the bearing capacity is taken as the stress when (sB) = 011 (or 11 foundation width)
minus0345 minus0345 05 s L 600 qmax = hs ∙ qtnet ∙ ( ) ∙ ( ) = 078 ∙ (9795) ∙ (011)05 ∙ ( )
B B 144
= 193 psi = 27860 psf qmax
Assuming a factor of safety (FS) of 3
qmax = 645 psi = 9287 psf
3
Step 6 Calculate settlement
68
119902119898119886119909 0345 2
1 frasl119865119878 L 119904 = 119861 ∙ [ ∙ ∙ ( ) ]
ℎ119904 119902119905119899119890119905 B
2 0345 1 645 psi 600 inches s = 144 inches ∙ [ ∙ ∙ ( ) ] = 18 inches
078 12311 psi 144 inches
Step 7 Determine if q gt qmax
q = 8000 psf lt 9287psf = qmax3
69
CHAPTER 4 DIRECT CPT METHOD FOR DEEP FOUNDATIONS
41 INTRODUCTION
The axial compression capacity (Qtotal) for a single pile includes a side component (Qside) end bearing
component (Qbase) and pile weight (Wpile) as shown in Figure 30 Since piles commonly push through
several layers a summation of the unit side frictions acting on the pile segments must be considered
over the length of the pile While the examples shown in this Guide resemble hand calculations
computer software is more efficient Programming the procedure is possible and represents a practical
method of designing deep foundations However commercial software is frequently available and is
MnDOTrsquos most common method of designing deep foundations
There are upwards of 40 different direct CPT methods that have been developed over the past five
decades to determine a piles axial compression capacity Many of the earliest methods relied on hand-
recorded information where qc data from mechanical CPTs would be collected at 20 cm intervals
whereas the direct CPT method uses scaled penetrometer readings via specified algorithms to obtain
the pile unit side friction and unit end bearing The method that will be used for deep foundation design
is the Modified UniCone method which uses all three readings of the electronic piezocone (qt fs and u2)
while addressing a variety of pile foundation types
Figure 30 Direct CPT evaluation of axial pile capacity
70
The modified UniCone Method is based upon a total of 330 pile load tests (three times the original
Unicone database) that were associated with SCPTu data Originally the UniCone method provided
approximate soil classification in five groups via a chart of qE vs fs (Figure 31) where qE = qt -u2 Later
using the modified approach with a larger data set provided soil sub classifications as shown in Figure
32 This new 9-zone normalized soil behavior type is determined using CPT data in combination with the
CPT Material Index
Figure 31 UniCone Method soil behavior type using CPT (Mayne 2017)
Figure 32 Modified UniCone Method soil behavior type using CPT (Mayne 2017)
71
42 MODIFIED UNICONE METHOD
In order to determine the axial pile capacity using the modified method the first step requires the
determination of geoparameters as shown in Direct CPT Method for Soil Characterization Once the
soil unit weight and CPT Material index are determined for each soil layer then the effective cone
resistance pile unit side friction (fp) and pile end bearing resistance (qb) can be determined
421 Step 1
Work through the steps provided in CPT Method for Soil Characterization until each soil layer is
defined by its CPT material index and soil behavior type using Figure 6 After these steps have been
completed continue to Step 2 to determine qE qb and fp
422 Step 2
Once qt is determined for each layer the effective cone resistance can be calculated using Equation 37
119902119864 = 119902119905 minus 1199062 37
Where (a) qE is the specific value at each elevation along the pile sides for determining fp and (b) at the bottom of the pile qE is averaged in the vicinity of the pile tip from the tip bearing elevation to about one diameter beneath the tip for determining qb
Using CPT material index and qE the pile end bearing resistance is calculated using Equation 38
38 119902119887 = 119902119864 ∙ 10(0325∙119868119888minus1218)
The pile unit side friction is obtained from qE and Ic
39 119891119901 = 119902119864 ∙ 120579119875119879 ∙ 120579119879119862 ∙ 120579119877119860119879119864 ∙ 10(0732∙119868119888minus3605)
Where θPT = coefficient for pile type (084 for bored 102 for jacked 113 for driven piles) θTC = coefficient for loading direction (111 for compression and 085 for tension) θRATE = rate coefficient applied to soils in SBT zone 1 through 7 (109 for constant rate of penetration test and 097 for maintained load tests)
72
423 Step 3
Due to piles extending through multiple layers the unit side components acting on various pile
segments would need to be summed if not using the direct CPT method Since CPT calculates data at
regular intervals of 2 cm to 5 cm along the sides of the pile the average fp in each layer can be used
directly in Equation 40 to obtain the shaft capacity
119876119904119894119889119890 = 119891119901 ∙ 119860119904 40
Where fp = average pile side friction along pile length from eqn 39 As = πmiddotdmiddotH d = pile diameter and H = length embedded below grade
The base capacity for a pile in compression loading is given by Equation 41 For piles in tension (or uplift)
Qbase can be taken as 0
= 119902119887 ∙ 119860119887 41 119876119887119886119904119890 Where
qb = end bearing resistance from eqn 38 Ab = πmiddotd24 (area of a circular pile)
424 Step 4
The final step is to calculate the axial pile capacity using Equation 42
119876119905119900119905119886119897 = 119876119904119894119889119890 + 119876119887119886119904119890 minus 119882119901119894119897119890 42
43 AXIAL PILE DISPLACEMENTS
Movement of pile foundations can be assessed using elastic continuum theory which has been
developed using finite element analyses boundary elements and analytical closed-form solutions In
the case of piles passing through several soil layers the elastic solution can be used by stacking pile
segments (each with its own stiffness) as represented by soil Youngrsquos modulus The use of software is
recommended for pile groups Several available programs such as DEFPIG GROUP and PIGLET can
handle pile groups under axial and lateralmoment loading
73
44 EXAMPLE PROBLEMS
441 Example 5 Direct CPT Methods Axial Pile Capacity
Several piles need to be placed beneath the edge of a building Use the CPT data collected for this site
shown in Figure 34 to determine the axial capacity for one of the piles Assume round steel driven piles
will be used with a diameter of 1275 inches wall thickness of 025 inches and lengths of 80 feet
Concrete will be used to fill the piles To solve for all geoparameters use the Direct CPT Method for Soil
Characterization section All sand layers can be assumed ldquodrainedrdquo with ν = 02
Figure 33 Deep foundation end bearing pile diagram
74
Figure 34 CPT data from Minnesota for example problem 5
75
Solution
Soil total unit weight
Estimate soil layering using ldquorules of thumbrdquo
Figure 35 Soil layers using ldquorules of thumbrdquo for pile capacity example using direct CPT method
Layer 1
From Figure 35 fs = 18 psi taken as a representative value of the layer
76
18 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1210 pcf
145 psi
Layer 2
From Figure 35 fs = 12 psi taken as a representative value of the layer
12psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1172 pcf
145 psi
Layer 3
From Figure 35 fs = 20 psi taken as a representative value of the layer
20 psi γt = 624 pcf ∙ [122 + 015 ∙ ln (100 ∙ + 001)] = 1219 pcf
145 psi
CPT Material Index
Layer 1
From Figure 35 fs = 18 psi and qc = 3000 psi
qt = qc + u2 ∙ (1 minus a) = 3000 psi + 20 psi ∙ (1 minus 08) = 3004 psi
∙ 4 feet = 1219 pcf ∙ 4 feet = 4877 psf = 34 psi σvo = γt
fs 18 psi Fr() = 100 ∙ = 100 ∙ = 060
(qt minus σvo) (3004 psi minus 34 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
77
uo = 0 psi
prime σvo = σvo minus uo = 34 psi minus 0 psi = 34 psi
(qt minus σvo)σatm (3004 psi minus 34 psi )145 psi = = Qtn prime )n (σvoσatm (34 psi145 psi)n
prime σvo 34 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(060)]2
Qtn = 3593 n = 038 lt 10 Ic = 14 (ie sand)
Layer 2
From Figure 35 fs = 12 psi and qc = 500 psi
qt = qc + u2 ∙ (1 minus a) = 500 psi + 40 psi ∙ (1 minus 08) = 508 psi
= 4838 psf + γt ∙ 45 feet = 4838 psf + 1172 pcf ∙ 45 feet = 5756 psf = 400 psi σvo
fs 12 psi Fr() = 100 ∙ = 100 ∙ = 260
(qt minus σvo) (508 psi minus 400 psi)
Step 2 Iterate to solve for Ic Steps are not shown for brevity
78
uo = 0 psi
prime σvo = σvo minus uo = 400 psi minus 0 psi = 400 psi
(qt minus σvo)σatm (508 psi minus 400 psi )145 psi = = Qtn prime (σvoσatm)n (400 psi145 psi)n
prime σvo 400 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(260)]2
Qtn = 117 n = 10 le 10 Ic = 29 (ie clayey silt)
Layer 3
From Figure 35 fs = 20 psi and qc = 5000 psi
qt = qc + u2 ∙ (1 minus a) = 5000 psi + 0 psi ∙ (1 minus 08) = 5000 psi
= 5756 psf + γt ∙ 6 feet = 5756 psf + 1219 pcf ∙ 6 feet = 6488 psf = 451 psi σvo
fs 20 psi Fr() = 100 ∙ = 100 ∙ = 040
(qt minus σvo) (5000 psi minus 451 psi)
79
Step 2 Iterate to solve for Ic Steps are not shown for brevity
prime σvo = σvo minus uo = 451 psi minus 0 psi = 451 psi
(qt minus σvo)σatm (5000 psi minus 451 psi )145 psi = = Qtn prime )n (σvoσatm (451 psi145 psi)n
prime σvo 451 psi n = 0381 ∙ Ic + 005 ( ) minus 015 = 0381 ∙ Ic + 005 ( ) minus 015
σatm 145 psi
Ic = radic[347 minus log(Qtn)]2 + [122 + log(040)]2
= 1801 n = 06 lt 10 = 15 (ie sand) Qtn Ic
Soil Behavior Type (SBT)
Layer 1
Based on values of Ic (14) Qtn (359) and Fr (06) the first layer is defined as a ldquoGravelly Sandrdquo
from Figure 36
80
Figure 36 Soil layer 1 using SBT method
81
Layer 2
Based on values of Ic (29) Qtn (117) and Fr (26) the second layer is defined as a ldquoSilty Mixrdquo
from Figure 37
Figure 37 Soil layer 2 using SBT method
82
Layer 3
Based on values of Ic (15) Qtn (180) and Fr (04) the third layer is defined as a ldquoSandrdquo from
Figure 38
Figure 38 Soil layer 3 using SBT method
Effective Cone Resistance
Layer 1
83
qE = qt minus u2 = 3004 psi minus 20 psi = 2984 psi
Layer 2
qE = qt minus u2 = 508 psi minus 40 psi = 468 psi
Layer 3
qE = qt minus u2 = 5000 psi minus 0 psi = 5000 psi
Pile End Bearing Resistance
Layer 3
= qE ∙ 10(0325∙Icminus1218) qb = 5000 psi ∙ 10(0325∙15minus1218) = 9085 psi
Pile Unit Side Friction
Layer 1
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 2984 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙14minus3605) = 99 psi
Layer 2
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 468 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙29minus3605) = 211 psi
Layer 3
minus3605) fp = qE ∙ θPT ∙ θTC ∙ θRATE ∙ 10(0732∙Ic
fp = 5000 psi ∙ 113 ∙ 111 ∙ 109 ∙ 10(0732∙15minus3605) = 202 psi
84
Axial Pile Capacity
The side capacity is based on pile design as shown in Figure 39
Layer 1
Qside = fp ∙ As = 99 psi ∙ (π ∙ d ∙ L) = 99 psi ∙ (π ∙ 1275 in ∙ 48 in) = 19064 lb
Layer 2
Qside = fp ∙ As = 211 psi ∙ (π ∙ d ∙ L) = 211 psi ∙ (π ∙ 1275 in ∙ 588 in) = 457122 lb
Layer 3
Qside = fp ∙ As = 202 psi ∙ (π ∙ d ∙ L) = 202 psi ∙ (π ∙ 1275 in ∙ 240 in) = 58170 lb
Figure 39 Soil layering compared to pilings
85
End Bearing Resistance in Layer 3
d2 1275 in2
Qbase = qb ∙ Ab = 9085 psi ∙ (π ∙ ) = 9085 psi ∙ (π ∙ ) = 115932 lb 4 4
Wp = (γsteel ∙ Asteel ∙ Depth) + (γsteel ∙ Aconc ∙ Depth)
Wp = (490 pcf ∙ 003 ft2 ∙ 55 ft) + (150 pcf ∙ 085 ft2 ∙ 55 ft) = 7724 lbs
Layer 3
Qtotal = ΣQside + Qblayer3 minus Wp
Qtotal = (19064 + 45713 + 58170) lbs + 115932 lbs minus 7724 lbs = 642563 lbs
= 642 kips Qtotal
Table 8 Summary of selected and calculated parameters
Layer L (in) qc
(psi)
fs
(psi)
Fr Ic Qtn qE
(psi)
qb
(psi)
Qbase
(lb)
fp
(psi)
Qside
(lb)
Qtotal
(kip)
1 48 3000 18 06 14 356 2984 NA NA 99 19064
2 588 500 12 26 29 117 468 NA NA 211 457122
3 240 5000 20 04 15 180 5000 9085 115932 202 58170
Total 115932 534355 642
86
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Agaiby SA (2018) Advancements in the interpretation of seismic piezocone tests in clays and other geomaterials PhD dissertation School of Civil amp Environmental Engineering Georgia Institute of Technology Atlanta GA
Agaiby SS and Mayne PW (2018) Interpretation of piezocone penetration and dissipation tests in sensitive Leda Clay at Gloucester Test Site Canadian Geotechnical Journal (accepted for publication 08 March 2018) CGJ-2017-0388
Amar S Baguelin F Canepa Y and Frank R (1998) New design rules for the bearing capacity of shallow foundations based on Menard PMT Geotechnical Site Characterization Vol 2 (Proc ISC-1 Atlanta) 727-733
American Petroleum Institute (API) (1981) Planning designing and constructing fixed offshore platforms Recommended practice API-RP2A 17th edition Washington DC API
American Petroleum Institute (API) (1987) Planning designing and constructing fixed offshore platforms Recommended practice API-RP2A 18th edition Washington DC API
American Petroleum Institute (API) (1989) Planning designing and constructing fixed offshore platforms Recommended practice API-RP2A 19th edition Washington DC API
Andersen KH and Stenhamar P (1982) Static plate loading tests on overconsolidated clay Journal of the Geotechnical Engineering Division ASCE 108 (GT7) 919-934
Basu D and Salgado R (2012) Load and resistance factor design of drilled shafts in sands Journal of Geotechnical amp Geoenvironmental Engineering 138 (12) 1455-1469
Berardi R Jamiolkowski M and Lancellotta R (1991) Settlement of shallow foundations in sands selection of stiffness on the basis of penetration resistance Geotechnical Engineering Congress Vol 1 185-200 (Proc GSP-27 Boulder) Reston Virginia ASCE
Brand EW Muktabhant C and Taechathummarak A (1972) Load tests on small foundations in soft clay Performance of Earth and Earth-Supported Structures Vol 1 Part 2 903-928 (Proc PC) ASCE Reston Virginia ASCE
Briaud JL and Gibbens RM (1999) Behavior of five large spread footings in sand Journal of Geotechnical amp Geoenvironmental Engineering 125 (9) 787-797
Briaud JL (2007) Spread footings in sand load-settlement curve approach Journal of Geotechnical amp Geoenvironmental Engineering 133 (8) 905-920
Brown DA Turner JP and Castelli RJ (2010) Drilled Shafts Construction Procedures and LRFD Design Methods Geotechnical Engineering Circular GEC 10 Report No FHWA NHI-10-016 Washington DC National Highway Institute US DOT
87
Burland J B (1973) Shaft Friction of Piles in Clay -A Simple Fundamental Approach Ground Eng 6(3) 30-42
Cai G Liu S and Puppala A (2012) Reliability assessment of CPTu-based pile capacity predictions in soft clay deposits Engineering Geology 84-91
Canadian Geotechnical Society (1992) Canadian Foundation Engineering Manual Third Edition Vancouver British Columbia BiTech Publishers
Cerato AB and Lutenegger AJ (2007) Scale effects of shallow foundation bearing capacity on granular material Journal of Geotechnical amp Geoenvironmental Engineering 133 (10) 1192-1201
Chen BS-Y and Mayne PW (1996) Statistical relationships between piezocone measurements and stress history of clays Canadian Geotechnical Journal 33 (3) 488-498
Cho W and Finno RJ (2010_ Stress-strain responses of block samples of compressible Chicago glacial clays Journal of Geotechnical amp Geoenvironmental Engineering 136 (1) 178-188
Chung SF Randolph MF and Schneider JA (2006) Effect of penetration rate on penetrometer resistance in clay Journal of Geotechnical amp Geoenvironmental Engineering 132 (9) 1188-1196
Clausen CJF Aas PM and Karlsrud K (2005) Bearing capacity of driven piles in sand the NGI approach Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis Group
Consoli NC Schnaid F and Milititsky J (1998) Interpretation of plate load tests on residual soil site Journal of Geotecnical amp Geoenvironmental Engingeering 124 (9) 857-867
Dasenbrock DD (2006) Assessment of pile capacity by static and dynamic methods to reconcile design predictions with observed performance Proc 54th Minnesota Geotechnical Engineering Conference 95-108 edited by Labuz J Univ of Minnesota St Paul
DeBeer E and Martens A (1957) Method of computation of an upper limit for the influence of heterogeneity of sand layers on the settlement of bridges 275-282 Proc ICSMFE-4 Vol 1 London ICSMFE
DeBeer EE (1965) Bearing capacity and settlement of shallow foundations on sand Proc Symposium on Bearing Capacity and Settlement of Foundations 15-33 Duke University Durham NC
Decourt L (1999) Behavior of foundations under working load conditions Proc PCSMGE-11 Foz do Iguassu Brazil Vol 4 453-487
Demers D and Leroueil S (2002) Evaluation of preconsolidation pressure and the overconsolidation ratio from piezocone tests of clay deposits in Quebec Canadian Geotechnical Journal 39 (1) 174-192
Eslaamizaad S and Robertson PK (1996) Cone penetration test to evaluate bearing capacity of foundations in sand 429-438 Proc CGCFG-49 Vol 1 St Johns Newfoundland
Eslami A Alimirzaei M Aflaki E and Molaabasi H (2017) Deltaic soil behavior classification using CPTu records Marine Georesources amp Geotechnology 35 (1) 62-79
88
Eslami A and Gholami M (2005) Bearing capacity analysis of shallow foundations from CPT data 1463-1466 Proc ICSMGE- 16 Vol 3 (Osaka) Millpress Rotterdam
Eslami A and Gholami M (2006) Analytical model for the ultimate bearing capacity of foundations from cone resistance Scientia Iranica 13 (3) 223-233
Eslami A and Fellenius BH (1997) Pile capacity by direct CPT and CPTu methods applied to 102 case histories Canadian Geotechnical Journal 34 (6) 886-904
Fellenius BH (2009) Basics of Foundation Design Vancouver BiTech Publishers
Fellenius BH and Altaee A (1994) Stress and settlement of footings in sand Vertical and Horizontal Deformations of Foundations and Embankments 2 1760-1773
Fellenius BH and Eslami A (2000) Soil profile interpreted from CPTu data Proc Geotechnical Engineering Conference Year 2000 Geotechnics Asian Inst of Technology Bangkok Thailand Available from wwwfelleniusnet
Finno R J and Chung C K (1992) Stress-strain-strength responses of compressible Chicago glacial clays Journal of Geotechnical Engineering 118 (10) 1607ndash1625
Fleming K Weltman A Randolph M and Elson K (2009) Piling Engineering Third Edition London Taylor amp Francis Group
Frank R and Magnan J-P (1995) Cone penetration testing in France national report Proceedings International Symposium on Cone Penetration Testing 3 (CPT95 Linkoumlping) Swedish Geotechnical Society Report 3(95) 147-156
Gavin K Adekunte A and OKelly B (2009) A field investigation of vertical footing response on sand Geotechnical Engineering 162 257-267
Hegazy YA (1998) Delineating geostratigraphy by cluster analysis of piezocone data PhD dissertation School of Civil amp Environmental Engineering Georgia Institute of Technology Atlanta GA
Holtz RD Kovacs WD and Sheehan TC (2011) An Introduction to Geotechnical Engineering 2nd Edition London Pearson Publishing
Hunt CE Pestana JM Bray JD and Riemer M (2002) Effect of pile driving on static and dynamic properties of soft clay Journal of Geotechnical amp Geoenvironmental Engineering 128 (1) 13-24
Jamiolkowski M Ladd CC Germaine JT and Lancellotta R (1985) New developments in field and lab testing of soils Proc ICSMFE-11 1 57-154
Jardine R Chow F Overy R and Standing J (2005) ICP design methods for driven piles in sands and clays London Thomas Telford Publishing
Jardine RJ Lehane BM Smith P and Gildea P (1995) Vertical loading experiments on rigid pad foundations at Bothkennar Geotechnique 45 (4) 573-597
89
Karlsrud K (2012) Prediction of load-displacement behavior and capacity of axially-loaded piles in clay based on analyses and interpretation of pile load test results PhD Dissertation Norwegian Univ Science amp Technology Trondheim Norway
Karlsrud K Clausen CJF and Aas PM (2005) Bearing capacity of driven piles in clay the NGI approach Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis
Karlsrud K Lunne T Kort DA and Strandvik S (2001) CPTU correlations for clays Proc 16th ICSMGE 2 683-702
Kimura T Kusakabe O and Saitoh K (1985) Geotechnical model tests of bearing capacity problems in a centrifuge Geotechnique 35 (1) 33-45
Knappett JA and Craig RF (2012) Craigs Soil Mechanics 8th Edition London Spon Press Taylor amp Francis
Kolk HJ and Velde EVD (1996) A reliable method to determine friction capacity of piles driven into clays Paper No 7993 Proc Offshore Technology Conference Houston
Kolk HJ Baaijens AE and Senders M (2005) Design criteria for pipe piles in silica sands Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis Group
Kulhawy FH (2004) On the axial behavior of drilled shafts Geosupport 2004 1-18
Kulhawy FH and Jackson CS (1989) Foundation Engineering Current Principles amp Practices 2 (GSP 22) Reston VA ASCE
Kulhawy FH and Mayne PW (1990) Manual for Estimating Soil Properties for Foundation Design EPRI Report EL-6800 Electric Power Research Institute Palo Alto 306 page Download from wwwepricom
Kulhawy FH Trautmann CH Beech JF ORourke TD and McGuire W (1983) Transmission Line Structure Foundations for Uplift-Compression Loading Report EL-2870 Palo Alto CA Electric Power Research Institute wwwepricom
Ku T and Mayne PW (2015) In-situ lateral stress coefficient (K0) from shear wave velocity measurements in soils Journal of Geotechnical amp Geoenvironmental Engineering 141 (12) 101061(ASCE) GT1943-56060001354
Ladd CC and DeGroot DJ (2003) Recommended practice for soft ground site characterization Soil amp Rock America 2003 3-57
Larsson R (1997) Investigations and Load Tests in Silty Soils SGI Report R-54 Linkoumlping Swedish Geotechnical Institute
Larsson R (2001) Investigations and Load Tests in Clay Till SGI Report R-59 Linkoumlping Swedish Geotechnical Institute
Lee J and Salgado R (2005) Estimation of bearing capacity of circular footings on sands based on cone penetration tests Journal of Geotechnical amp Geoenvironmental Engineering 131 (4) 442-452
90
Lehane BM (2003) Vertically loaded shallow foundation on soft clayey silt Geotechnical Engineering 156 (1) Thomas Telford London 17-26
Lehane BM (2013) Relating foundation capacity in sands to CPT qc Geotechnical amp Geophysical Site Characterization 1 London Taylor amp Francis Group
Lehane BM Doherty JP and Schneider JA (2008) Settlement prediction for footings on sand Deformational Characteristics of Geomaterials (IS-Atlanta) 1 Rotterdam Millpress
Lehane BM Li Y and Williams R (2013) Shaft capacity of displacement piles in clay using the CPT Journal of Geotechnical amp Geoenvironmental Engineering 139 (2) 253-266
Lehane BM Schneider JA and Xu X (2005) The UWA-05 method for prediction of axial capacity of driven piles in sand Frontiers in Offshore Geotechnics (Proc ISFOG Perth) London Taylor amp Francis
Low HE Lunne T Andersen KH Sjursen MA Li X and Randolph MF 2010 Estimation of intact and remoulded undrained shear strengths from penetration tests in soft clays Geotechnique 60 (11) 843-859
Lunne T Robertson PK and Powell JJM (1997) Cone Penetration Testing in Geotechnical Practice London RoutledgeBlackie Academic
Lutenegger AJ and DeGroot DJ (1995) Settlement of Shallow Foundations on Granular Soils Report Contract 6332 Task Order 4 submitted to Massachusetts Highway Dept by University of Mass-Amherst Amherst MA
Lutenegger AJ and Adams MT (2003) Characteristic load-settlement behavior of shallow foundations Proc FONDSUP 2 381-392
Mase T and Hashiguchi K (2009) Numerical analysis of footing settlement problem by subloading surface model Soils amp Foundations 49 (2) 207-220
Mayne PW (1987) Determining preconsolidation stress and penetration pore pressures from DMT contact pressures Geotechnical Testing Journal 10(3) 146-150
Mayne PW (1991) Determination of OCR in Clays by Piezocone Tests Using Cavity Expansion and Critical State Concepts Soils and Foundations 31 (2) 65-76
Mayne PW (2007) NCHRP Synthesis 368 on Cone Penetration Test Washington DC Transportation Research Board National Academies Press wwwtrborg
Mayne PW (2009) Geoengineering Design Using the Cone Penetration Test Richmond BC Canada ConeTec Inc
Mayne PW (2014) Keynote Interpretation of geotechnical parameters from seismic piezocone tests Proceedings 3rd Intl Symp on Cone Penetration Testing 47-73
Mayne PW (2016) Evaluating effective stress parameters and undrained shear strengths of soft-firm clays from CPT and DMT Australian Geomechanics Journal 51 (4) 27-55
91
Mayne PW (2017) Stress history of soils from cone penetration tests 34th Manual Rocha Lecture Soils amp Rocks Vol 40 (3) Saouml Paulo ABMS
Mayne PW Christopher B Berg R and DeJong J (2002) Subsurface Investigations -Geotechnical Site Characterization Publication No FHWA-NHI-01-031 Washington DC National Highway Institute Federal Highway Administration
Mayne PW Coop MR Springman S Huang A-B and Zornberg J (2009) Geomaterial Behavior and Testing Proc 17th Intl Conf Soil Mechanics amp Geotechnical Engineering 4 Rotterdam MillpressIOS Press
Mayne PW and Illingworth F (2010) Direct CPT method for footings on sand using a database approach Proc ISCPT-2 3 315-322
Mayne PW and Niazi F (2017) Recent developments and applications in field investigations for deep foundations Proceedings 3rd Bolivian Conference on Deep Foundations 1 141-160
Mayne PW Peuchen J and Baltoukas D (2015) Piezocone evaluation of undrained strength in soft to firm offshore clays Frontiers in Offshore Geotechnics III 2 1091-1096
Mayne PW and Poulos HG (1999) Approximate displacement influence factors for elastic shallow foundation systems ASCE Journal of Geotechnical amp Geoenvironmental Engineering 125 (6) 453-460
Mayne PW Uzielli M and Illingworth F (2012) Shallow footing response on sands using a direct method based on cone penetration tests Full Scale Testing and Foundation Design (GSP 227) Reston Virginia ASCE
Mayne PW and Woeller DJ (2014) Direct CPT method for footings on sands silts and clays Geocharacterization for Modeling and Sustainability (GSP 234) 1983-1997
McClelland B (1974) Design of Deep Penetration Piles for Ocean Structures Journal Geot Eng Div 100(GT7) 705-747
Mesri G (1975) Discussion on new design procedure for stability of soft clays ASCE Journal of the Geotechnical Engineering Division 101(GT4) pp 409-412
Meyerhof GG (1956) Penetration tests and bearing capacity of cohesionless soils Journal of Soil Mechanics amp Foundations Division 82 (SM1) 1-19
Meyerhof GG (1974) General Report Penetration testing outside Europe Proceedings ESOPT-1 Vol 21 Swedish Geotechnical Society Stockholm
Meyerhof GG (1976) Bearing capacity and settlement of pile foundations Journal of Geotechnical Engineering Division 102(3) 197-228
Missiaen T Verhegge J Heirman K and Crobe P (2015) Potential of cone penetrating testing for mapping deeply buried palaeolandscapes in the context of archaeological surveys in polder areas Journal of Archaeological Science 55 174-187
92
Mlynarek Z Wierzbicki J Goglik S and Bogucki M (2014) Shear strength and deformation parameters of peat and gyttja from CPTu DMT and VST Proceedings 5th International Workshop on CPTu and DMT in soft clays and organic soils 193-210 Poznan Poland Polish Committee on Geotechnics Menard Polska Tensar Internationa
Nejaim PF Jannuzzi GMF and Danziger FAB (2016) Soil behavior type of the Sarapuiacute II test site Geotechnical amp Geophysical Site Characterization 5 1009-1014
Niazi FS and Mayne PW (2013) Cone penetration test based direct methods for evaluating static axial capacity of single piles Geotechnical and Geological Engineering 31 (4) 979-1009
Niazi FS and Mayne PW (2015) Enhanced Unicone expressions for axial pile capacity evaluation from piezocone tests Proc International Foundations Conference amp Equipment Expo RestonVA ASCE
Niazi FS and Mayne PW (2016) CPTu-based enhanced UniCone method for pile capacity Engineering Geology 212 21-34
Nowacki F Karlsrud K and Sparrevik P (1996) Comparison of recent tests on OC clay and implications for design Large Scale Pile Tests in Clay London Thomas Telford
ONeill MW (2001) Side resistance in piles and drilled shafts Journal of Geotechnical amp Geoenvironmental Engineering 127 (1) 3-16
Ouyang Z and Mayne PW (2018) Effective friction angle of clays and silts from piezocone Canadian Geotechnical Journal in press doiorg101139cgj-2017-0451
Paikowsky SG Canniff MC Lesny K Kisse A Amatya S and Muganga R (2010) LRFD Design and Construction of Shallow Foundations for Highway Bridge Structures NCHRP Report 651 Washington DC National Cooperative Highway Research Program Transportation Research Board
Pestana JM Hunt CE and Bray JD (2002) Soil deformation and excess pore pressure filed around a closed-ended pile Journal of Geotechnical amp Geoenvironmental Engineering 128 (1) 1-12
Poulos HG and Davis EH (1974) Elastic Solutions for Soil and Rock Mechanics New York Wiley amp Sons
Poulos HG and Davis E (1980) Pile Foundation Analysis amp Design New York John Wiley amp Sons
Powell JJM and Lunne T (2005) Use of CPTU data in clays and fine-grained soils Studia Geotechnica et Mechanica XXVII(3) 29-66
Randolph MF (2003) Rankine Lecture Science and empiricism in pile foundation design Geotechnique 53 (10) 847-875
Robertson PK (1990 1991) Soil classification using the cone penetration test Canadian Geotechnical Journal 27 (1) 151-158 Closure 28 (1) 176-178
Robertson PK (1991) Estimation of foundation settlements in sand from CPT Geotechnical Engineering Congress 2 Reston Virginia ASCE
93
Robertson PK (2009) Cone penetration testing a unified approach Canadian Geotechnical J 46 (11) 1337-1355
Robertson PK (2009a) Performance based earthquake design using the CPT Keynote Lecture International Conference on Performance-Based Design in Earthquake Geotechnical Engineering IS Tokyo Tsukuba Japan
Robertson PK (2009b) Interpretation of cone penetration tests a unified approach Canadian Geotechnical Journal 46 (11) 1337-1355
Robertson PK and Cabal KL (2007) Guide to cone penetration testing Second Edition Signal Hill California Gregg In-Situ
Robertson PK and Cabal KL (2015) Guide to Cone Penetration Testing for Geotechnical Engineering 6th Edition Signal Hill California Gregg Drilling amp Testing Inc
Schmertmann JH (1970) Static cone to compute static settlement over sand Journal of the Soil Mechanics and Foundations Division 95 (SM3) 1011-1043
Schmertmann JH (1978) Guidelines for cone penetration test performance and design Report FHWA-TS-78-209 Washington DC Federal Highway Administration
Schnaid F Wood WR Smith AKC and Jubb P (1993) Investigation of bearing capacity and settlements of soft clay deposits at Shellhaven Predictive Soil Mechanics London Thomas Telford
Schneider JA Xu X and Lehane BM (2008) Database assessment of CPT-based design methods for axial capacity of driven piles in siliceous sands J Geotechnical and Geoenvironmental Engineering 134 (9) 1227-1244
Semple D (1980) Recent Developments in the Design amp Construction of Piles London Institution of Civil Engineers
Senneset K Sandven R amp Janbu N (1989) Evaluation of soil parameters from piezocone tests In-Situ Testing of Soil Properties for Transportation Transportation Research Record 1235 24-37
Shahri AA Melehmir A and Juhlin C (2015) Soil classification analysis based on piezocone penetration test data Engineering Geology 189 32-47
Stuedlein AW and Holtz RD (2010) Undrained displacement behavior of spread footings in clay The Art of Foundation Engineering Practice GSP 198 Reston Virginia ASCE
Takesue K Sasao H and Matsumoto T (1998) Correlation between ultimate pile skin friction and CPT data Geotechnical Site Characterization 2 1177ndash1182
Tand KE Funegard EG and Briaud J-L (1986) Bearing capacity of footings on clay CPT method Use of In-Situ Tests in Geotechnical Engineering GSP 6 1017-1033
Tand KE Warden PE and Funegaringrd EG (1995) Predicted-measured bearing capacity of shallow footings on sand PISCPT 2 589-594
Thomas D (1968) Deep soundings test results and the settlement of spread footings on normally consolidated sands Geotechnique 18 (2) 472-488
94
Tomlinson MJ (1957) The adhesion of piles driven into clay soils Proc 4th Intl Conf on Soil Mechanics and Foundation Engineering 2 66-71
Tomlinson MJ (1986) Foundation Design amp Construction London Longman Scientific
Tomlinson MJ (1987) Pile Design and Construction Practice London Viewpoint Publication
Trofimenkov JG (1974) Penetration testing in the USSR Proceedings ESOPT-1 1 147-154
Tumay MT Hatipkarasulu Y Marx ER and Cotton B (2013) CPTPCPT-based organic material profiling Proc 18th Intl Conf on Soil Mechanics amp Geotechnical Engineering Paris 633-636
Uzielli M and Mayne PW (2011) Statistical characterization and stochastic simulation of load-displacement behavior of shallow footings Proc Georisk 2011 Geotechnical Risk Management amp Assessment (GSP 224) 672-679
Uzielli M and Mayne PW (2012) Load-displacement uncertainty of vertically-loaded shallow footings on sands and effects on probabilistic settlement estimation Georisk Assessment and Management of Risk for Engineered Systems and GeoHazards 6 (1) 50-69
Valsson SM (2016) Detecting quick clay with CPTu Proceedings 17th Nordic Geotechnical Meeting Reykajavik Iceland Challenges in Nordic Geotechnics
Van Dijk BFJ and Kolk HJ (2011) CPT-based design method for axial capacity of offshore piles in clay Frontiers in Offshore Geotechnics II 555-560
Vesić AS (1975) Bearing capacity of shallow foundations Foundation Engineering Handbook New York Van Nostrand Reinhold Publishers
Vesic AS (1977) NCHRP Synthesis of Highway Practice 42 Design of Pile Foundations Washington DC Transportation Research Board
Yu F and Yang J (2012) Base capacity of open-ended steel pipe piles in sand J Geotechnical and Geoenvironmental Engineering 138 (9) 1116-1128
Zawrzykraj P Rydelek P and Bakowska A (2017) Geoengineering properties of Eemian peats from Radsymin (central Poland) in the light of static cone penetration and dilatometer tests Engineering Geology 226 290-300
95
APPENDIX A
List of Figures
Figure A1 Conventional methods versus direct CPT approach to shallow foundation response A-3
Figure A2 Direct bearing capacity relationship for embedded square and strip footings situated on clean
sands (after Schmertmann 1978) A-6
Figure A3 Direct CPT bearing factor Rk as function of footing width B and embedment depth from finite
element analyses (Tand et al 1995) A-6
Figure A4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio and sand
consistency (Eslaamizaad amp Robertson 1996) A-7
Figure A5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp Salgado (2005) in
terms of footing size relative density and base settlement A-8
Figure A6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and normalized
cone tip resistance (after Eslami amp Gholami 2005) A-8
Figure A7 Direct relationship between ultimate bearing stress in clays and measured cone tip resistance
(Schmertmann 1978) A-10
Figure A8 Direct relationship between ultimate foundation bearing stress in clays and net cone tip
resistance (after Tand et al 1986) A-11
Figure A9 Measured load-displacements for each of the five large footings at TAMU (Briaud amp Gibbens
1994) sand site A-12
Figure A10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU (Briaud amp
Gibbens 1994) footings A-13
Figure A11 Characteristic stress versus square root normalized displacements for TAMU footings A-15
Figure A12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sand A-16
Figure A13 Summary of sand formation factors with corresponding CPT resistances A-19
Figure A14 Normalized foundation stresses vs square root of normalized displacement for spread
footings on sand (modified after Mayne et al 2012) A-20
Figure A15 Definitions of capacity from three types of observed foundation load test responses
(modified after Kulhawy 2004) A-21
Figure A16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footings A-22
A-1
Figure A17 Measured load vs displacement curves for 3 shallow foundations on clay at Baytown Texas
(Stuedlein amp Holtz 2010) A-25
Figure A18 Characteristic stress vs square root of normalized displacement for all three foundations at
Baytown site (Stuedlein amp Holtz 2010) A-26
Figure A19 Normalized foundation stress vs square root of normalized displacements A-27
Figure A20 Applied foundation stress vs CPT-calculated stress for 67 footings A-28
Figure A21 Applied footing stress vs CPT calculated stress on logarithmic scales A-29
Figure A22 Summary graph and equations of direct CPT method for evaluating the vertical A-30
Figure A23 Foundation soil formation parameter hs versus CPT material index Ic A-32
Figure A24 Bearing capacity ratio qmaxqnet versus CPT material index Ic A-33
Figure A25 Influence factor for rectangular foundations from elastic theory solution (Mayne amp
Dasenbrock 2017) A-34
List of Tables
Table A1 Direct CPT methods for bearing capacity of footings on clean sands A-4
Table A2 Direct CPT methods for bearing capacity of footings on clays A-9
Table A3 Summary of large footings soil conditions and reference sources of database A-17
Table A4 Sources for shallow foundation performace A-34
A-2
A1 INTERPRETATION OF SOIL PARAMETERS FROM CONE
PENETRATION TESTS
A11 Introduction
Geotechnical site characterization is an important first step towards the evaluation of
subsurface conditions and determination of soil layering geomaterial classification and the
evaluation of soil engineering parameters for the analysis and design of foundations retaining
walls tunnels excavations embankments and slope stability Increasing use of the electronic
cone penetrometer in highway applications is prominent in the USA because the results are
obtained much faster and less expensive than traditional methods that rely on rotary drilling
augering sampling and laboratory testing Moreover the cone penetration test (CPT) provides
at least three independent and continuous readings on soil behavior that are digitally recorded
and fully available at the end of the sounding As such the reliable and consistent
interpretation of CPT data is important since civil engineering works will best realize the
efficiency and economy of this technology in constructed facilities for county state and
interstate projects
A111 Cone Penetration Testing
The cone penetrometer is an electronically-instrumented steel probe that is vertically pushed into the
ground at constant rate of 20 mms (08 insec) using a hydraulic system The penetrometer is outfitted
with load cells pressure transducers and inclinometers that measure at least three readings
continuously with depth (a) cone tip resistance qt (b) sleeve friction fs and (c) porewater pressure u2
With the latter reading the device is called a piezocone thus the designation CPTu is used to indicate
porewater pressure Additional sensors are available that can be used to measure inclination
resistivity pH temperature shear wave velocity and other readings if desired
Figure A1 shows a schematic rendering of the cone penetration test that is performed in the
field per ASTM D 5778 procedures The hydraulic system is nominally rated at 180 kN (20 tons)
capacity and often mounted on a truck but can also be positioned on tracked vehicles or
anchored frames Figure A2 shows a MnDOT cone truck that uses the full dead weight of the
rig for the hydraulic reaction forces which are applied at the center of the vehicle The cab is
enclosed so that soundings may proceed during inclement weather and the data acquisition
system and operator are protected
Figure A 1 Setup and procedures for co ne penetration testing (CPT)
A-4
Figure A2 CPT truck-mounted rig used by MnDOT
A representative sounding taken at the I-35 test site just northeast of Saint Paul is presented in Figure
A3 Here the separate profiles of qt fs and u2 with depth are shown in side-by-side plots In the last
graph the hydrostatic porewater pressure (uo) due to the groundwater table is indicated by the blue
dashed line
These readings are used to interpret the soil layers types and properties of the ground as
discussed in the following sections
A-5
Figure A3 Representative CPT sounding at I-35 test site northeast of St Paul MN showing profiles of
(a) cone resistance (b) sleeve friction and (c) penetration porewater pressures
A112 Soil Parameter Interpretation
A variety of soil engineering parameters can be interpreted from CPT results on the basis of
theoretical frameworks analytical models or numerical simulations otherwise by empirical
methods based on correlations and statistics of databases Figure A4 shows a conceptual
utilization of the readings from the CPT for interpretation of geostratigraphy compressibility
flow characteristics and stress-strain-strength behavior of soils
Table A1 lists a selection of geoparameters that have been addressed for these purposes The
most common or useful relationships will be discussed in subsequent sections and reference is
made to other sources (Lunne et al 1997 Mayne 2007 Robertson amp Cabal 2015) for additional
details
A-6
Figure A4 Conceptual post-processing of CPT results for geoparameter evaluation
Table A1 List of common geoparameters interpreted from CPT data
Symbol Parameter Remarks Notes
SBT Soil behavior type (SBT)
Ic CPT material index
γt Unit weight
ρt Mass Density
σvo Overburden stress
u0 Hydrostatic pressure
A-7
σvo Effective stress
Table A1 Continued
Symbol Parameter Remarks Notes
σp Preconsolidation stress σp = 033 (qnet)m where m = fctn (Ic)
(or Effective Yield Stress) where stresses are in kPa
YSR Yield stress ratio YSR = σpσvo (formerly OCR)
c Effective cohesion intercept
ϕ Effective friction angle
su Undrained shear strength
D Constrained modulus
G0 Small-strain modulus
G Shear modulus
E Youngs modulus
cvh Coefficient of consolidation
K Bulk modulus
A-8
LI
K0 Lateral stress coefficient
k Hydraulic conductivity
Liquefaction index
A12 Geostratigraphy and Soil Behavioral Type (SBT)
In routine CPT soil samples are not normally taken and thus the measured stress friction andor
porewater pressure readings are used to infer the types of soils This can be accomplished using
three basic approaches
a Rules of Thumb or approximate guidelines for a quick visual assessment
b Soil Behavioral Type (SBT) Charts that are based on either the three readings (qt fs u2) net readings
including net cone resistance (qnet = (qt-σvo) effective cone resistance (qE = qt - u2) friction ratio (Rf =
100middotfsqt) and excess porewater pressures or using normalized CPT readings such as Q F Bq or U
c Probabilistic methods where the CPT readings have been calibrated from lab tests on
recovered soil samples (eg Tumay et al 2013)
A121 Approximate Rules of Thumb
The approximate rules of thumb provide simple guidelines for soil type and suggest that sands are
identified when qt gt 5 MPa and u2 asymp u0 whereas intact clays are prevalent when qt lt 5 MPa and u2 gt u0
(Mayne et al 2002) The magnitude of porewater pressures reflects the consistency of the intact clay
such that soft (u2 asymp 2middotu0) firm (u2 asymp 4middotu0) stiff (u2 asymp 8middotu0) and hard (u2 asymp 20middotu0) However for fissured
overconsolidated clays the measured porewater pressures are less than hydrostatic in fact often
negative u2 lt 0 On land negative porewater pressure readings at the shoulder filter element of the
piezocone should be greater than -1 bar ie u2 gt -100 kPa (-147 psi) Also for clean sands the friction
ratio (FR = 100middotfsqt lt 1) while for insensitive clays (FR gt 4)
A-9
Figure A5 presents a 36-m deep piezocone sounding from the MnDOT Wakota Bridge site which crosses
the Mississippi River at I-494 (Dasenbrock 2006) The qt and u2 readings have been annotated using the
aforementioned rules of thumb indicating the predominance of sands at this site As noted there are
five interbedded clay layers evident at depths of 1 m 3-4 m 7 - 12 m 17 m and 22-27 m
Figure A5 Interpretation of soil types from CPT data taken from the Wakota Bridge on I-494 using
approximate rules of thumb
A122 Soil Behavioral Type Charts
The most common approach for identification of soil types is based on soil behavioral charts
Reviews of several chart methods are provided elsewhere (Kulhawy amp Mayne 1990 Fellenius amp
Eslami 2000 Shahri et al 2015) Current popularity favors the 9-zone classification system to
evaluate soil behavioral type (SBT) that uses normalized piezocone readings (Robertson 1990
1991 Lunne et al 1997)
A-10
(119902119905 minus 120590119907119900) (a) normalized tip resistance 119876 = A11 prime 120590119907119900
100∙119891119904 (b) normalized sleeve friction 119865() = A12
(119902119905 minus 120590119907119900)
(1199062 minus 1199060) (c) normalized porewater pressure 119861119902 = A13
(119902119905 minus 120590119907119900)
While the Q-F-Bq classification is a three-dimensional plot it is often presented as a set of two-dimensional
plots specifically Q vs F and Q vs Bq as shown in Figure A6 Data are grouped according to layers and
can be superimposed to identify their association with the nine soil behavioral types All indirect CPT soil
classification approaches should be cross-checked and verified for a particular geologic setting and local
geotechnical conditions before routine use in practice
Figure A6 Nine-zone soil behavioral type charts (a) normalized cone resistance vs normalized
sleeve friction (b) normalized cone resistance vs normalized porewater pressure parameters
(adapted after Robertson 1991 Lunne et al 1997)
The development of a CPT material index (Ic) has been found advantageous in the initial screening of soil
types and helps to organize the sounding into their respective SBT zones of similar soil response In this
case the CPT index is found from (Robertson 2009)
A-11
A2 22 )log221()log473( rtnc FQI
where Qtn = modfied stress-normalized net cone resistance given by
(119902119905 minus 120590119907119900)120590119886119905119898 119876 = A31 prime (120590119907119900120590119886119905119898)119899
where σatm = reference stress equal to atmospheric pressure (1 atm asymp 1 bar = 100 kPa asymp 1 tsf asymp 147 psi)
The exponent n is soil-type dependent n = 1 (clays) n asymp 075 (silts) and n asymp 05 (sands) If units of bars are
used then Qtn (units of bars) is simply
(119902119905 minus 120590119907119900) 119876 = 119899 A32
prime ) (120590119907119900
The operational value of exponent n is found by iteration Initially an exponent n = 1 is used to calculate
the starting value of Ic (ie Qtn = Q) and then the exponent is upgraded to
prime 120590119907119900 119899 = 0381 ∙ 119868119888 + 005 ( ) minus 015 le 10 A4 120590119886119905119898
Then the index Ic is recalculated Iteration converges quickly and generally only 3 cycles are needed to
secure the operational Ic at each depth The soil zones and associated Ic values are detailed in Figure A7
The sensitive soils of zone 1 can be screened using the following expression
A-12
Zone 1 119876119905119899 lt 12119890119909119901( minus14 ∙ 119865119903)
A5
If any soil layers are found within zone 1 (sensitive soils) then caution should be exercised as these
structured clays are prone to instability collapses and difficulties in construction and performance
Another indicator of zone 1 sensitive soils is when the normalized porewater parameter Bq gt 08 When
these criteria are evident the geoengineer should consult with senior geotechnical personnel or the chief
engineer for guidance
Figure A7 Delineation of soil behavioral type zones using CPT material index Ic
A-13
Stiff overconsolidated clayey sands and sandy clays soils of zone 8 (15 lt Fr lt 45) and zone 9 (Fr gt
45) can be identified from the following criterion
Zones 8 and 9 0020)1(00030)1(0050
12
rr
tnFF
Q A6
The remaining soil types are identified by the CPT material index Zone 2 (organic clayey soils Ic ge 360)
Zone 3 (clays to silty clays 295 le Ic lt 360) Zone 4 (silt mixtures 260 le Ic lt 295) Zone 5 (sand mixtures
205 le Ic lt 260) Zone 6 (clean sands 131 le Ic lt 205) and Zone 7 (gravelly to dense sands Ic le 131)
The red dashed line at Ic = 260 represents an approximate boundary separating drained (Ic lt 260) from
undrained behavior (Ic gt 260)
Once the specific zone has been assigned to a layer a visual representation can be made to show
either the zone number or colorization so that the predominant layers and soil types can be
realized
A123 Probabilistic Soil Types from CPT
The inference of soil type has been addressed using probabilistic methods as discussed by Abu-
Farshakh et al (2008) and Tumay et al (2013) Here the CPT readings can be post-processed to
provide the percentage of sand silt and clay components with depth The output is consistent
and compatible with the current MnDOT soil classification chart (Figure A8) which is similar in
content to that used by the USDA
A-14
Figure A8 Chart for soil classification system by MnDOT
A13 Identifying Soft Organic Soils
The identification of soil type using cone penetration tests (CPT) is usually done on the basis of empirical
soil behavioral type (SBT) charts that use the measured readings (qt fs andor u2) While there at least
25 different sets of charts available (ie Hegazy 1998 Eslami et al 2017 Valsson 2016) one of the most
widely used and popular is the 9-zone SBT system that uses three normalized cone readings Qtn Fr and
Bq (Robertson amp Cabal 2007 Lunne et al 1997) The system is denoted as SBTn and plotted on charts in
terms of Q vs F and Q vs Bq as shown in Figure A9
A-15
1
10
100
1000
01 1 10
Norm
aliz
ed
Con
e T
ip R
esis
tan
ce
Q
Normalized Friction Fr = 100 fst(qt - svo) ()
9-Zone Soil Behavioral Type Chart
Gravelly Sands(zone 7)
Sands
(zone 6)
Sandy Mixtures(zone 5)
Silt Mix(zone 4)
Clays
(zone 3)
Organic Soils(zone 2)
Sensitive Clays
and Silts(zone 1)
Very stiff OC clayto silt (zone 9)
Very stiffOC sandto clayey
sand(zone 8)
1
10
100
1000
-06 -04 -02 00 02 04 06 08 10 12 14
No
rmal
ized
Co
ne
Tip
Res
ista
nce
Q
Normalized Porewater Bq = Du2(qt-svo)
Clays(zone 3)
Sensitive(zone 1)Organic
(zone 2)
Gravelly Sands(zone 7)
Sands(zone 6)
Figure A9 CPT charts for SBTn system (a) Q versus F (b) Q versus Bq
Of particular concern in geotechnical site characterization is the proper identification of soft
organic soils such as organic clays and silts (OL and OH) and peats (Pt) These soils often cause
difficulties and problems during construction and performance of civil engineering works because
of bio-degradation long-term creep excessive settlements andor other issues
The SBTn system has a category labelled Zone 2 recognizes organic soils However several recent
studies have noted that data from CPTu soundings do not always fall within the bounds
delineated using Zone 2 even though laboratory tests and field inspections clearly note the
presence of organic soils Lab tests to classify organic soils includes loss on ignition (LOI gt 5)
and two pairs of Atterberg limits testings per ASTM standards as well as the visual-manual
identification of strong organic odor and smell and dark color notably black and dark gray
Results by Mlynarek et al (2014) show that organic soils in Poland are not adequately identified by the
Q-F chart as seen in Figure A10 Moreover studies by Zawrzykraj et al (2017) found that neither the Q-
A-16
F and Q-Bq plots worked well to recognize organic soils and peats in Poland Nejaim et al (2016) found
that Brazilian soft organic clays were misclassified as Zone 3 (clays) rather than Zone 2 (organic clays)
Similarly a recent study on CPTu data from organic clays located at 24 sites in the USA Sweden Mexico
Brazil and Australia have been compiled (Agaiby 2018) These data are shown to generally avoid falling
with the Zone 2 boundaries in either Q-F or Q-Bq charts thus are not recognized during these soundings
as indicated by Figure A11 The exception in this case is the Mexico City clay which is correctly identified
however the other 23 sites are generally not properly recognized and categorized Additional findings by
Missiaen et al (2015) found that the CPTu results in Belgian soils also did not identify organic clays and
peats in the non-normalized version of these charts that uses Qt versus friction ratio (Rf = 100middotfsqt) as
detailed by Robertson amp Cabal (2007)
Figure A10 Soil behavioral chart with superimposed CPTu data from Polish organic soils
(from Mlynarek et al 2014)
A-17
Figure A11 Paired SBTn charts with CPTu data from 24 organic clay sites indicating non-compliance
with Zone 2 (organic soils)
(compiled by Agaiby 2018)
A14 Simplified Yield Stress Evaluation in Clays by CPTu
From the development of a hybrid formulation of spherical cavity expansion (SCE) theory with critical-
state soil mechanics (CSSM) the effective preconsolidation stress or yield stress (σp) of clays can be
expressed in terms of three piezocone parameters (a) net cone tip resistance qnet = qt - σvo (b)
measured excess porewater pressure Δu2 = u2 - u0 and (c) effective cone resistance qE = qt - u2 Details
on the SCE-CSSM solutions are given elsewhere (Mayne 1991 Mayne 2017 Agaiby amp Mayne 2018)
A-18
For normal and well-behaved clays which include inorganic fine-grained soils such as clays and silts
of low sensitivity a set of simple relationships can be derived By adopting characteristic default values
for effective friction angle ϕ = 30deg and rigidity index (IR = 100) linear expressions for the effective
preconsolidation stress are obtained
prime 120590119901 = A7 033 ∙ 119902119899119890119905
prime 120590119901 = 053 ∙ ∆1199062 A8
prime 120590119901 = 060 ∙ 119902119864 A9
A141 Application to soft Chicago clay
An example of a common geomaterial in this category includes the infamous soft Chicago clays
that were deposited in a glacio-lacustrine environment (Cho amp Finno 2010) The soft clay has a
mean water content of 20 liquid limit of 38 and plasticity index PI= 12 Results of CPTu tests
conducted at the national geotechnical experimentation site (NGES) at Northwestern University
are shown in Figure A12 (Ouyang amp Mayne 2017) As seen in the profile a thin soft clay resides
between depths of 9 to 105 m with a thicker soft clay layer in the range of 12 to 18 m
A-19
0
2
4
6
8
10
12
14
16
18
20
00 02 04 06 08 10 12D
ep
th (
m)
CPTu qt and u2 (kPa)
sand (SP)
Blodgett
Park Ridge
Deerfieldsoft clay
ClayLayerof Study
Northwestern UnivNational Test Site
Figure A12 Results of CPTu sounding at the NGES at Northwestern University Illinois
Post-processing of the CPTu data using Equations A7 A8 and A9 show good agreement in evaluating the
profile of effective yield stress in this clay layer compared with results from one-dimensional
consolidation tests made on undisturbed samples taken at the site
A-20
10
11
12
13
14
15
16
17
18
19
20
0 100 200 300 400 500 600 700 800 900 1000
De
pth
(m
)Effective Yield Stress sp (kPa)
Consols
033 qnet
053 Du2
060 qE
svo
Figure A13 Evaluation of effective yield stresses at NWU using consolidation tests and CPTu
A142 Application to soft Bay Mud California
Another example of a well-behaved fine-grained soil is the San Francisco Bay Mud which is a soft clay
deposit in northern California Results of a representative CPTu are shown in Figure A14 and indicate the
soil profile at a site in eastern San Francisco (Hunt et al 2002) For this site index values include 70 lt
wn lt 90 60 lt LL lt 80 and 35 lt PI lt 45 Post-processing the CPTu data using Equations A7 A8 and
A9 show good agreement with each other as well as with the results of consolidation tests as seen in
A-21
Figure A15 The consolidation tests were performed using a CRS device as reported by Pestana et al
(2002)
San Francisco Bay Mud (Pestana et al 2002)
0
5
10
15
20
25
30
35
40
0 1000 2000 3000
De
pth
(m
ete
rs)
Cone Resistance qt (kPa)
0
5
10
15
20
25
30
35
40
0 10 20 30 40 50 60
Sleeve Friction fs (kPa)
0
5
10
15
20
25
30
35
40
0 500 1000 1500
Porewater u2 (kPa)
Miscellaneous Fill
Young Bay Mud
Young Bay Mud
Clayey Sand
Miscellaneous Fill
Young Bay Mud
Young Bay Mud
Figure A14 Representative CPTu in San Francisco Bay Mud (data from Hunt et al 2002)
A-22
San Francisco Bay Mud (Pestana et al 2002)
0
5
10
15
20
25
30
0 100 200 300 400 500 600
De
pth
(m
ete
rs)
Preconsolidation Stress sp (kPa)
svo
033 qnet
053 Du2
060 qE
CRS
svo
Figure A15 Evaluation of effective yield stresses in Bay Mud from consolidation tests and CPTu
A15 CPTu Screening of Soft Organic Clays
For soft organic clays the Equations A7 A8 and A8 will not apply thus can serve as a means to
screen such geomaterials from the soil profile Two examples of CPTu soundings in soft organic
clays are presented to illustrate the approach Additional case records and documentation of
CPTu data in organic clays are given in Agaiby (2018)
A151 Soft organic alluvial soils in Washington DC
Results of in-situ tests including CPTu soundings in soft organic clayey silts along the Potomac River at the
Anacostia Naval Air Station are presented by Mayne (1987) Here the soft soils are categorized as organic
clayey silts (OH) per the Unified Soils Classification Systems (USCS) with mean values (and plus and minus
A-23
one standard deviations) of wn = 68 plusmn 16 LL = 83 plusmn 25 and PI = 37 plusmn 17 A representative
piezocone sounding is depicted in Figure A16 For the post-processing of CPTu data the use of the Q-F
and Q-Bq graphs do not find any points within the Zone 2 category for organic soils When the data are
evaluated to determine the profile of effective yield stress the Equations A7 A8 and A9 do not agree as
shown by Figure A17
0
5
10
15
20
25
0 2000 4000 6000
Dep
th (
met
ers)
Cone Resistance qt (kPa)
0
5
10
15
20
25
0 20 40 60 80 100
Sleeve Friction fs (kPa)
0
5
10
15
20
25
0 200 400 600
Pore Pressure u2 (kPa)
Figure A16 Representative CPTu in soft organic clay along the Potomac River DC
A-24
0
5
10
15
20
25
0 3 6 9D
epth
(m
eter
s)Soil Behaviour Type Zone
Q and Fchart
0
5
10
15
20
25
0 100 200 300 400 500 600
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
033qnet
053 Du2
06 qE
Oed
Figure A17 Post-processing of CPTu data from Anacostia-Bolling (a) SBTn zones using CPT material
index Ic (b) CPTu screening equations for yield stress
Note that the hierarchy of the results from Equations A7 A8 and A9 show that the CPTu
evaluations for preconsolidation stress in soft organic clays indicates
[A10] σp = 053 Δu2 lt 033 qnet lt 060 qE
A152 Soft organic peaty clay in Saint Paul MN
Results of CPTu tests were collected during a training exercise underneath I-35E just northeast
of Saint Paul MN in 2007 All three MnDOT CPT rigs available at the time were used to obtain
in-situ test data at the site The site is well known as having soft organic soils and samples were
obtained from three borings made at the site (Boring ID T523 T542 and T518) Boring logs
A-25
indicate the presence of highly organic silt loams with marl and spongy organic clayey silts
From 12 samples the natural water contents ranged from 30 to 191 with a mean of 112 plusmn
58
A representative piezocone sounding from the site (MnDOT CPTu ID F22Y0703C) is used for
this illustration and is presented as Figure A18 Post-processing these data using the SBTn
algorithms and CPT material index show that the soils classify primarily as Zone 3 (clays) and
Zone 4 (silty mix) thus do not identify the geomaterials properly as Zone 2 (organic soils)
0
5
10
15
0 500 1000 1500 2000
Dep
th (
met
ers)
Cone Tip Resistance
qt (kPa)
0
5
10
15
0 10 20 30 40 50 60
Sleeve Friction
fs (kPa)
0
5
10
15
-100 0 100 200 300 400
Porewater Pressure
u2 (kPa)
Figure A18 MnDOT CPTu sounding at the I-35E test site near St Paul Minnesota
A-26
1
10
100
1000
01 1 10
Norm
aliz
ed
Con
e T
ip R
esis
tan
ce
Q
Normalized Friction Fr = 100 fst(qt - svo) ()
9-Zone Soil Behavioral Type Chart
St PaulGravelly Sands(zone 7)
Sands
(zone 6)
Sandy Mixtures(zone 5)
Silt Mix(zone 4)
Clays
(zone 3)
Organic Soils(zone 2)
Sensitive Clays
and Silts(zone 1)
Very stiff OC clayto silt (zone 9)
Very stiffOC sandto clayey
sand(zone 8)
1
10
100
1000
-06 -04 -02 00 02 04 06 08 10 12 14
No
rmal
ized
Co
ne
Tip
Res
ista
nce
Q
Normalized Porewater Bq = Du2(qt-svo)
Clays(zone 3)
Sensitive(zone 1)Organic
(zone 2)
Gravelly Sands(zone 7)
Sands(zone 6)
Figure A19 Post-processing St Paul sounding using the 9-zone SBTn classification system
Using the recommended approach Figure A20 shows the CPT-evaluated yield stress profiles from
equations A7 A8 and A9 Here again the three estimates of σp do not agree but show the same
hierarchy as detailed in Equation A10
A-27
0
5
10
15
20
0 50 100 150 200 250 300 350 400
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
06 qeff svo
033qnet Consols
054 Du2
Figure A20 CPTu screening of soft organic soils at St Paul test site
A16 CPTu Evaluations of Yield Stress in Soft Organic Soils
Once the presence of soft organic clays and silts is identified using the aforementioned CPT
screening algorithms the evaluation of effective yield stress is conducted using the procedure
outlined in Chapter 2 of this MnDOT manual For soft organic soils an exponent of m = 090 is
recommended (Mayne et al 2009) Therefore the preconsolidation stress is obtained from
(units of kPa)
prime = A11 120590119901 033 ∙ (119902119899119890119905)090
The algorithm can be expressed in dimensionless form by
prime = A12 120590119901 033 ∙ (119902119905 minus 120590119907119900)119898prime ∙ (120590119886119905119898100)1minus119898prime
A-28
where σatm = reference stress (= 100 kPa = 147 psi) and m = 090 for soft organic silts and clays
The approach is applied to the two former example case studies in Figure A21 (Washington DC)
and Figure A22 (St Paul MN) respectively showing good agreement with benchmark results
taken from laboratory consolidation tests on undisturbed samples of these sites in both cases
0
5
10
15
20
25
0 100 200 300 400 500 600
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
033qnet^090
Oed
Figure A21 Profiles of yield stress from consolidation tests and CPTu method for organic clayey silts at
Anacostia-Bolling site in Washington DC
A-29
0
5
10
15
20
0 20 40 60 80 100 120 140 160
Dep
th (
met
ers)
Effective Yield Stress sp (kPa)
svo
CPT
Consols
120590119901prime = 033 119902119899119890119905
090 in kPa
Figure A22 Profiles of yield stress from consolidation tests and CPTu method for organic clayey silts at
MnDOT test site near Saint Paul Minnesota
A17 Soil Unit Weight
As shown by Figure A23 total soil unit weight (t) can be estimated from the CPT sleeve friction resistance
(Mayne 2014) The equation can be expressed by
120574119905 = 120574119908 ∙ [122 + 015 ∙ 119897119899 (100 lowast 119891119904120590119886119905119898 + 001)] A13
where w = unit weight of water and satm = atmospheric pressure
A-30
Figure A23 Evaluation of soil unit weight from CPT sleeve friction
A18 Effective Stress Friction Angle
Sands
The effective stress friction angle () is one of the most important soil properties as it governs the strength
of geomaterials as well as affects soil-pile interface and pile side friction While an effective cohesion
intercept (c) can also be considered this is usually reserved for cemented or bonded geomaterials or
unsaturated soils and may lose its magnitude with time ageing or with prolonged environmental
changes
For clean quartz to silica sands where porewater pressures are essentially hydrostatic (Bq = 0) the
following expression has been calibrated with triaxial compression test results from undisturbed sand
samples and normalized cone resistances as presented in Figure A24 (Mayne 2007 2014)
A-31
120601prime(119889119890119892) = 176deg + 110deg 119897119900119892 (1199021199051) A14
Figure A24 Evaluation of effective stress friction angle in quartz-silica sands from CPT
Note Relationship applies to drained soil behavior when Ic lt 26 andor Bq lt 01
where qt1 is an earlier form of stress normalized cone tip resistance for sands given by
(119902119905 120590119886119905119898) 1199021199051 = A151 05 prime (120590119907119900120590119886119905119898)
A-32
In terms of units of bars this is expressed simply as (in units of bars)
119902119905 1199021199051 = 05 A152 prime ) (120590119907119900
Recently Robertson amp Cabal (2015) recommended the use of the modified form of normalized cone
resistance (Qtn) in Equation A10
120601prime (119889119890119892) = 176deg + 110deg 119897119900119892 (119876119905119899) A16
In sands qnet asymp qt since the overburden term is small relative to the cone tip resistance Figure A25 shows
that the two normalizations give very comparable results
Figure A25 Comparable values from qt1 and Qtn normalization schemes for CPTs in sands
Clays and Silts
A-33
In the case of soft to firm intact clays and silty clays the effective friction angles are determined from the
normalized cone resistance and porewater pressure parameters (Senneset et al 1989 Mayne 2016) as
shown in Figure A26 The exact solution when the angle of plastification = 0 is given as
qBQ
)tan1(tan61
1)tanexp()245(tan 2
A17
Approximate NTH Solution for from CPTu
qBQ
)tan1(tan61
1)tanexp()245(tan2
Approx asymp 295deg∙Bq0121∙[0256 + 0336∙Bq + log Q ]
1
10
100
20 25 30 35 40 45
Co
ne
Res
ista
nce
Nu
mb
er Q
(degrees)
Bq
010204060810
c = 0 = 0deg
Theory (dots)
Approximation (lines)
Theory
Figure A26 Evaluation of effective stress friction angle in clays and silts from CPTu results
Note Relationship applies to undrained soil behavior when Ic ge 26 andor Bq ge 01
which can be approximately inverted into the form (Mayne 2007)
A-34
0121 120601prime = 295deg ∙ 119861119902 ∙ [ 0256 + 0336 ∙ 119861119902 + log(119876)] A18
This algorithm is specifically applicable for the following ranges of porewater pressure parameter (01 le
Bq lt 1) and effective stress friction angles (20deg le lt 45deg)
A19 Stress History
The stress history can be characterized by an apparent yield stress or preconsolidation stress (sp) as well
as by its normalized and dimensionless form YSR = spsvo = yield stress ratio The YSR is in effect the
same as overconsolidation ratio (OCR) however now generalized to accommodate mechanisms of
preconsolidation that occur beyond just erosion glaciation and removal of overburden stresses but also
due to ageing desiccation repeated cycles of wetting-drying bonding repeated freeze-thaw cycles
groundwater changes and other factors
A-35
A generalized approach for evaluating the yield stress or preconsolidation in soils using net cone
resistance has been formulated as presented in Figure A27 (Mayne 2015) Yield stress in soils from CPT
10
100
1000
10000
10 100 1000 10000 100000
Yie
ld S
tre
ss
s
y
(kP
a)
Net Cone Resistance qt - svo (kPa)
Blessington
General Trend
sy = 033(qt-svo)m
Fissured Clays m = 11
Intact clays m = 10 Sensitive clays m = 09
Silt Mixtures m = 085Silty Sands m = 080
Clean Sands m = 072
m = 11 10 09
085
080
072
Units kPa
Fissured Clays
Figure A27 Evaluation of yield stress or preconsolidation stress in soils from CPT
The algorithm can be expressed in dimensionless form by
1minus119898prime prime 120590119886119905119898 120590119901 = 033 ∙ (119902119905 minus 120590119907119900)119898prime
( ) A19 100
where m = exponent depends on soil type m = 1 (intact clays) 085 (silts) 080 (sandy silts to silty sands)
and 072 (sands) For fissured clays the exponent m may be 11 or higher depending upon the age of the
A-36
formation and degree of jointing and discontinuities Note that fissured clays can be identified when Ic lt
26 and Bq lt 01
The exponent m has also been calibrated with CPT material index as presented in Figure A28
An algorithm to express this relationship for non-fissured soils is given by
25)652(1
2801
cIm
A20
This allows the post-processing of CPT to automatically choose the appropriate exponent m for
a layer by layer analysis
Figure A28 Yield stress exponent m in terms of CPT material index Ic
A-37
A110 Lateral Stress Coefficient
The horizontal geostatic state of stress is represented by the lateral stress coefficient K0 = shosvo
commonly referred to as the at-rest condition The magnitude of K0 for soils that have been loaded and
unloaded can be approximately estimated from
119870119900 = (1 minus 119904119894119899120601prime) ∙ 119874119862119877119904119894119899120601prime A21
Data compiled from in-situ K0 measurements using self-boring pressuremeter tests (SBPMT) total stress
cells (TSC) or push-in spade cells andor laboratory methods (instrumented consolidometers triaxials
andor suction measurements) on a variety of clays silts and sands have been compiled and reported by
Ku amp Mayne (2015) as shown in Figure A29 verifying Equation A15 as a means for evaluating K0 in soils
Figure A29 Relationship between lateral stress coefficient K0 and YSR or OCR in soils (Ku amp Mayne
2015)
A-38
A111 Undrained Shear Strength
The loading of soils can occur under conditions of being fully drained (Du = 0) partially drained or full
undrained (DVV0) where Du = excess porewater pressures (above hydrostatic) and DVV0 = volumetric
strain Further details on specific stress paths are best explained in terms of critical state soil mechanics
or CSSM (Mayne et al 2009 Holtz Kovacs amp Sheahan 2011) The prevailing drainage conditions depend
upon the rate of loading and permeability characteristics of the soil Normally in sands that are pervious
and exhibit high permeability a drained response occurs Exception to this may occur in loose sands
during fast earthquake loading resulting in soil liquefaction In clays that exhibit low permeability a fast
rate of loading will result in undrained loading at constant volume This in fact is a temporary and
transient condition often termed short term loading In the long term eventually porewater pressures
will dissipate (albeit slowly) and a drained response will prevail thus termed long term loading
The overall soil strength is controlled by the effective stress strength envelope Most commonly this is
represented by a simple linear relationship termed the Mohr-Coulomb criterion where the maximum
shear stress (max called the shear strength) is given as
= 119888prime + 120590prime ∙ 119905119886119899 120601prime A22 120591119898119886119909
where c = effective cohesion intercept s = effective normal stress and = effective stress friction angle
As a starting point values of c = 0 and = 30deg can be adopted for all soil types at least until the specific
soil type geologic formation and results from high-quality laboratory or field test data are available
Peak Undrained Shear Strength from Stress History
Using simplified critical state soil mechanics the peak undrained shear strength (max = su) can be
evaluated from the effective stress strength envelope (c = 0 effective friction angle ) and stress history
(ie OCR) in the form
prime DSS 119904119906 = (119904119894119899120601prime2) ∙ (119874119862119877)Λ ∙ 120590119907119900 A23
A-39
which corresponds to a simple shear mode Figure A30 presents the aforementioned relationship
together with data from 17 different clays tested in simple shear
Figure A30 Normalized undrained shear strength from simple shear tests on various clays
versus YSR or OCR
For the triaxial compression (TC) mode the equation would give slightly higher strengths that are
calculated from
prime TC 119904119906 = (1198721198882) ∙ (1198741198621198772)Λ ∙ 120590119907119900 A24
where Mc = (6 middot sin)(3-sin)
Peak Undrained Shear Strength from CPTu
A more direct approach to assessing undrained shear strength is via bearing capacity theory
whereby
A-40
A25 kt
votu
N
qs
s
where Nkt = bearing factor that depends on mode of shearing sensitivity of the clay degree of
overconsolidation and other factors For soft offshore clays the back figured Nkt from data collected at
14 well-documented sites (Low et al 2010) determined mean values based on mode of shearing Nkt =
119 (triaxial compression) Nkt = 136 (simple shear) and Nkt = 133 (vane)
A recent study of 51 clays that were tested by both field piezocone and laboratory CAUC triaxial tests
showed that essentially Nkt = 12 for intact clays and clayey silts of low to medium sensitivity (Mayne
Peuchen amp Baltoukas 2015) Figure A31 shows the slightly different trends for offshore versus onshore
clays For sensitive clays a lower value of Nkt = 10 would be appropriate and for fissured clays a higher
value of Nkt would be in the range of 20 to 30
A-41
A-42
Figure A31 Database from 51 clays with CPT qnet versus lab measured triaxial shear strength
(Mayne Peuchen amp Baltoukas 2015)
For soft to firm clays an alternate means to evaluate undrained shear strength is via the excess porewater
pressures ( u = u2 - u0) from the following
u
uN
us
D
D 2
A26
where N u = porewater bearing factor also dependent on the aforementioned factors The study by
Low et al (2010) found N u = 59 (triaxial compression) N u = 69 (simple shear) and N u = 71 (vane
shear)
A third alternate is found from the effective cone resistance (qE = qt - u2) whereby
kE
tu
N
uqs 2
A27
where NkE is a bearing term for this approach Mayne et al (2015) indicated a mean value of NkE = 8 for
soft-firm intact clays
Remolded Undrained Shear Strength from CPT
The remolded undrained shear strength (sur) is obtained from either field vane tests lab mini-vane shear
tests or lab fall cone devices This affords the evaluation of the clay sensitivity (St) which is defined as the
ratio of peak to remolded strengths at a given water content
119904119906(119901119890119886119896) 119878119905 = frasl A28 119904119906119903
For the CPT the sleeve friction has been noted to give values that are comparable to the
remolded undrained shear strength (Powell amp Lunne 2015) Therefore
asymp 119891119904 A29 119904119906119903
A112 Ground Stiffness and Soil Moduli
The stiffness of the ground can be represented by several geoparameters including the compressibility
indices (Cr Cc Cs) spring constants (kv) and moduli Regarding the latter there are several moduli that
are used in geotechnical engineering including shear modulus (G and Gu) Youngs modulus (E and Eu)
constrained modulus (D) bulk modulus (K) resilient modulus (MR) and subgrade reaction modulus (ks)
A-43
Soil Modulus
The definition of modulus is taken as E = DsD with Figure A32 showing a typical deviator stress (q = s1
- s3) versus axial strain (a) curve Theoretical interrelationships between the elastic moduli G E D and
K have a dependence on the Poissons ratio v The value of Poissons ratio can be taken as vu = 05 for
undrained loading (ie constant volume) while for drained loading which is accompanied by volumetric
strains a value of v = 02 may be used If we adopt the reference modulus as E then the
interrelationships with the other elastic moduli are given by
a Reference Stiffness 119864prime = drained Youngs modulus A30
119864prime
b Shear Modulus 119866prime = A31 [2(1+120584prime)]
119864prime (1minus120584prime) c Constrained Modulus 119863prime = A32
[(1+120584prime)(1minus2120584prime)]
119864prime
d Bulk Modulus 119870prime = A33 [3middot(1minus2120584prime)]
A-44
Figure A32 Representative stress-strain-strength curve for soils in triaxial compression
For the extreme case when = 0 in fact D = E When = 02 the two moduli are only 10
different D = 11middotE thus the constrained modulus and drained Youngs modulus are often
considered somewhat interchangeable
The resilient modulus (MR) applies to pavement analysis and design most commonly measured by cyclic
triaxial testing under repeated load applications In fact MR is a special case of Youngs modulus that
relates to the small strain stiffness measured in the nondestructive range but has developed permanent
plastic strains after many cycles of loading (Brown 1996 Dehler amp Labuz 2007)
The subgrade reaction modulus is actually a combined soil-structural parameter as its value
depends on the ground stiffness and the size of the loaded element The subgrade modulus is
defined as ks = q where q = applied stress and = measured deflection In terms of elasticity
solutions the deflection of a flexible circle of diameter d is given by = qmiddotdmiddot(1-2)E Therefore
119864prime
119896119904 = A34 [119889middot(1minus1205842)]
which has units of kNm3 or pcf
Table A2 summarizes several selected studies towards the approximate evaluation of these moduli
obtained directly from measured CPT data Note however that the results from in-situ penetrometer
tests generally represent a peak strength as indicated in Figure A32 Thus the measured cone tip
resistance (qt) reflects the top of the stress-strain curve either the undrained strength in clays (su) or the
effective friction angle () of sands or even a strength intermediate between these two values Thus
A-45
Source Soils Studied Reference Modulus
Findings
Kulhawy amp Mayne 1990
8 stiff OC clays Lab constrained modulus
D asymp 825 middot (qt - svo)
Mohammed et al (2000)
Resilient modulus
Abu-Farsakh 2004
7 Louisiana sites
Lab constrained modulus
D asymp 358 middot (qt - svo)
Mayne (2007b)
All soil types Constrained moduli from lab consolidation
First-order estimate D asymp middot(qt - svo)
Robertson (2009)
All soil types Constrained modulus from consolidation
D = m middot (qt - svo)
when Ic gt 22
1 m = Qtn when Qtn le 14
2 m = 14 when Qtn gt 14
when Ic lt 22 then
= 003 middot 10055 Ic + 168 m
Liu et al (2016)
16 clay sites in China
Resilient modulus MR = (146qt 053 + 1355 fs
14 + 236)244
(all in MPa)
Casey et al (2016)
8 clays Triaxial tests for undrained Youngs modulus
Eu(svc)07 = 316 - 23 middot LL
where Eu and svc (MPa)
and LL = liquid limit ()
qt is probably not the best means to obtain the slope of the stress-strain curve Instead the SCPTu that
provides the initial tangent shear modulus (Gmax) from the shear wave velocity (Vs) is a better choice
Table A2 Selected Modulus Relationships with Cone Penetration Test Measurements
A-46
Nonlinear Modulus
The small-strain shear modulus (Gmax or G0) represents the initial stiffness of all soils and rocks It is the
beginning of all stress-strain-strength curves for geomaterials and obtained from elasticity theory
119866119898119886119909 = 1198660 = 120588119905 middot 1198811199042 A35
where Vs = shear wave velocity t = tga = total soil mass density t = total soil unit weight and ga =
gravitational acceleration constant (98 ms2)
From a general viewpoint on stiffness the shear modulus G of soil is defined as the slope of shear stress
versus shear strain G = DDs for a tangent definition and by G = s as a secant definition The shear
modulus is related to its associated Youngs modulus E = 2G(1+v) Both moduli are in fact highly
nonlinear ranging from a maximum value at the small-strain stiffness (Gmax = G0 = t middotVs2 where t =
total soil mass density and Vs = shear wave velocity) to intermediate G values at medium strains (asymp 1)
to low values at peak strength As such a variety of algorithms and formulae have been developed to
represent either a partial range or the full stress-strain-strength behavior of soils over a range of
interests (eg Mayne 2005) In these formulations a variable number of input parameters may be
required in order to produce a stress-strain-strength curve
A modified hyperbolic form suggested by Fahey amp Carter (1993) has favorable attributes in that the
modulus reduction factor can be established with only a single variable This allows the initial stiffness
to the be small-strain stiffness (G0) that is reduced in terms of level of mobilized strength eg max =
qqmax = 1FS which is simply the reciprocal of the calculated factor of safety (FS) The magnitude of
secant shear modulus (G) corresponding to the particular level of loading is given by
119866 = 119872119877119865 middot 1198660 = (1198661198660) middot 1198660 A36
A-47
where the MRF = modulus reduction factor determined as
1 minus (120591 )119892 119872119877119865 = (1198661198660) = frasl A37 120591119898119886119909
and g = empirical exponent term Figure A33 shows the relationships for normalized shear stress ()
versus shear strain (s) and corresponding normalized secant shear modulus (G = s) with shear strain
over a range of exponent values 01 le g le 10 For a discussion of tangent G please see Fahey amp Carter
(1993)
Using laboratory data from resonant column-torsional shear tests and triaxial specimens with local
strain measurements for Gmax reference values modulus reduction curves for a selection of sands and
clays are presented in Figure A34 The data indicate that the exponent g falls within the range 02 lt g lt
05 for many soils tested drained and undrained A mean value of g = 03 is recommended for
preliminary site investigations and designs until additional information can be obtained
The value of max is the shear strength of the soil generally taken as either (1) drained (c = 0) max =
svo middot tan or (2) undrained where max = su = undrained shear strength as discussed earlier Thus each
shear stress () can be associated with its shear modulus (G) and the relevant shear strain is found from
s = G This allows for generation of nonlinear stress-strain-strength curves at all depths from SCPTu
data in clays sands and mixed soil types
A-48
Modified Hyperbola (Fahey amp Carter 1993 Canadian Geotech J) Coefficient Term f = 1
``
0
01
02
03
04
05
06
07
08
09
1
000001 00001 0001 001 01 1
No
rmal
ized
GG
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
0
01
02
03
04
05
06
07
08
09
1
0 01 02 03 04 05 06 07 08 09 1
No
rmali
zed
GG
ma
x
Mobilized Shear Stress max
g =1
07
05
03
02
01
0
01
02
03
04
05
06
07
08
09
1
00001 0001 001 01 1
No
rmali
zed
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
0
01
02
03
04
05
06
07
08
09
1
0 1 2 3 4 5 6 7 8 9 10
No
rmali
zed
max
Shear Strain s ()
g = 1
g = 07
g = 05
g = 03
g = 02
g = 01
Reduction Factor
GGmax = 1- (max)g
Figure A33 Normalized modulus (GGmax) and normalized shear stress (max) versus shear strain (s)
for modified hyperbolic relationship of Fahey amp Carter (1993)
A-49
0
01
02
03
04
05
06
07
08
09
1
0 01 02 03 04 05 06 07 08 09 1
Mo
du
lus
Re
du
cti
on
G
Gm
ax
or
EE
ma
x
Mobilized Strength max or qqmax
NC SLB Sand
OC SLB Sand
Hamaoka Sand (e=0625)
Hamaoka Sand
Toyoura Sand e = 067
Toyoura Sand e = 083
Ham River Sand
Ticino Sand
Kentucky Clayey Sand
Kaolin
Fujinomori Clay
Pietrafitta Clay
London Clay
Vallericca Clay
Pisa Clay
Pisa Clay
Pisa Clay
Pisa Clay
Kiyohoro Silty Clay
Thanet Clay
Exp g = 02
Exp g = 03
Exp g = 04
Exp g = 05
MRF = 1 - (qqmax)g
Figure A34 Modulus Reduction Factors (MRF) with mobilized strength of clays and sands
A113 Coefficient of Consolidation
The rate at which foundation and embankment settlements occur as well as the dissipation of excess
porewater pressures is controlled by the coefficient of consolidation (cv) The magnitude of cv is also
required in the design of vertical wick drains that can be installed in soft ground to expedite the time for
consolidation Using the results of CPTu dissipation tests which measure the rate at which the u2 readings
vary with time the in-situ profile of cv can be evaluated Often the piezocone-interpreted values of cv
are validated by comparison with results from laboratory one-dimensional consolidation tests (eg Abu-
Farsakh MY 2004) Alternatively a better method is to cross-check the values with the measured full-
scale performance of instrumented embankments that are constructed over the ground and the recorded
time-rate of consolidation and settlements can provide the best cv for that site (Abu-Farsakh MY et al
2011)
A-50
Monotonic Dissipation Tests
An illustrative dissipation curve from piezocone testing at IDT State Highway 95 at an embankment and
bridge crossing is presented in Figure A35 After the CPTu sounding was halted at a depth of 512 feet
the recorded decay of porewater pressures was observed to be monotonic with time until the dissipations
were ended at 1000 seconds During that time the u2 readings decreased from 115 psi to 47 psi At this
location the depth to the groundwater table is about 16 feet thus the calculated equilibrium water
pressure is 15 psi Commonly a characteristic time for piezo-dissipations is taken at 50 degree of
consolidation although other degrees may be adopted By evaluating the value of u2 at 50 as shown in
Figure A35 the characteristic t50 = 404 s can be obtained
0
20
40
60
80
100
120
1 10 100 1000 10000
Pore
wat
er P
ress
ure
u2
(psi
)
Time (s)
CPTU-02E-1 z = 1560 m
uo (hydrostatic)
u2 (initial) = 115 psi
u2 (final projected) = u0 = 15 psi
u2 at 50 = frac12(115+15) = 65 psi
Idaho DOT State Route 95Dissipation at 512 feet
t50 = 404 seconds
Time to reach50 consolidation
A-51
Figure A35 Piezo-dissipation at Sandpoint Bridge Crossing Idaho for depth z = 512 feet illustrating the
evaluation of characteristic time t50
It is also common to plot normalized excess porewater pressures relative to the measured initial Du2
that is obtained during penetration at the constant rate of 20 mms ie DuDui versus time Figure A36
shows a set of piezo-dissipation records for a bridge and embankment site in southern Louisiana
reported by the LTRC While the porewater pressure axis is arithmetic the time scale is often plotted on
either logarithmic or square root scales In either case the t50 is then found from the measured
dissipation curve when DuDui = 050
A-52
Figure A36 Set of monotonic piezo-dissipations taken at various depths for Courtableau Bridge
Site on Louisana State Highway 103 (Abu-Farsakh et al 2011)
Dilatory Dissipation Curves
In a number of situations a monotonic type dissipation does not occur on the onset but instead a dilatory
type curve is observed Here after stopping the push of the penetrometer the recorded porewater
pressures initially begin to rise and eventually reach a peak value then afterwards follow a phase where
the Du values decrease to hydrostatic (Figure A37) A full solution is available for both cases (Burns amp
Mayne 2002) Dilatory responses are most often associated with overconsolidated clays and silts
although occasionally are observed in fine-grained soils with low OCRs
The selection of the characteristic t50 value may found by using a square root of time plot to find the initial
u2 value by projection of the post-peak dissipatory data back to the ordinate axis as shown by the example
in Figure A38 In this example the initial u2 = 480 kPa and then the same procedures in Figure A35may
apply
A-53
Figure A37 Monotonic and dilatory porewater dissipation responses in soils
(after Burns amp Mayne 1998)
Figure A38 Method for obtaining t50 from dilatory dissipation curves
(Schneider amp Hotstream 2010)
A-54
Interpretation of Dissipation Test Data
There are a number of different solutions available for the interpretation of the coefficient of
consolidation (cv) from the piezocone dissipation curves For monotonic type response the strain path
method (SPM) developed at Oxford University (Teh amp Houlsby 1991) is well-recognized while the Georgia
Tech solution by Burns amp Mayne (2002) is based on spherical cavity expansion and critical state soil
mechanics (SCE-CSSM) and handles both monotonic and dilatory curves For both solutions a simplified
procedure can be recommended that relies on the aforementioned t50 value obtained from the measured
field data as given in Table A3 The units for cv are in cm2s as shown or equivalent
Table A3 Recommended procedures for calculating cv from t50 obtained in dissipation tests
Method Equation for coefficient of
consolidation cv
RemarksNotes Eqn
No
SPM
(Teh amp
Houlsby
1991) 50
2)(2450
t
Iac
Rc
v
t50 = measured time to reach 50
dissipation
IR = Gsu = undrained rigidity index
G = shear modulus
su = undrained shear strength
ac = penetrometer radius (ac = 178
cm for 10-cm2 cone ac = 220 for a
15-cm2 size)
A32
SCE-CSSM
(Burns amp
Mayne 2002) 50
7502 )()(0300
t
Iac Rc
v
A33
A-55
Evaluating rigidity index
Both the SPM and SCE-CSSM solutions require an estimate of the in-situ rigidity index (IR = Gsu) of the
soil If the results of SCPTU are available Krage et al (2014) have derived an expression for IR that
depends upon the small-strain shear modulus net cone tip resistance and effective overburden stress
250750
max50
8111
vonet
Rq
GI
s A38
where consistent units are input for Gmax qnet and svo The value of Gmax is obtained from B29 using the
measured shear wave velocity (Vs) and unit weight evaluated from B7
If the shear wave velocity is not measured it can be estimated from the CPT data A number of
methods have been reviewed for CALTRANS by Wair et al (2012) including one that is applicable
to sands silts and clays of low sensitivity that are inorganic and uncemented
119881 (119898119904) = [101 middot 119897119900119892(119902119905) minus 114]167 middot (100 middot 119891119904119902119905)03 A39 119904
where qt is in units of kPa (Hegazy amp Mayne 1995) An alternative relationship is given by Robertson amp
Cabal (2015)
A114 Hydraulic Conductivity
The hydraulic conductivity is a parameter that expresses the flow characteristics of the soil In
geotechnics it is also called the coefficient of permeability (k) and has units of cms or feetday
Through consolidation theory the hydraulic conductivity relates directly to the coefficient of
consolidation
A-56
A40 D
ck wv
where w = unit weight of water and D = constrained modulus
Therefore one approach to evaluating k can be from the site-specific cv obtained from the
dissipation tests using an estimate of D from one of the relationships given in Table A2
An approach developed for soft normally-consolidated soils that uses t50 directly to assess the magnitude
of k is presented in
Figure A39 (Parez and Fauriel 1988) The mean trendline through the wedges of soil type can be
approximated by
251
50 (sec)251
1)(
tscmkh
A41
A-57
Figure A39 Hydraulic conductivity versus dissipation time for 50 consolidation
(Parez amp Fauriel 1988)
A-58
APPENDIX B
List of Figures
Figure B1 Conventional methods vs direct CPT approach to shallow foundation response B-3
Figure B2 Direct bearing capacity relationship for embedded square and strip footings situated on clean
sands (after Schmertmann 1978) B-6
Figure B3 Direct CPT bearing factor Rk as function of footing width B and embedment depth from finite
element analyses (Tand et al 1995) B-6
Figure B4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio and sand
consistency (Eslaamizaad amp Robertson 1996) B-7
Figure B5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp Salgado (2005) in
terms of footing size relative density and base settlement B-8
Figure B6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and normalized
cone tip resistance (after Eslami amp Gholami 2005) B-8
Figure B7 Direct relationship between ultimate bearing stress in clays and measured cone tip resistance
(Schmertmann 1978) B-10
Figure B8 Direct relationship between ultimate foundation bearing stress in clays and net cone tip
resistance (after Tand et al 1986) B-11
Figure B9 Measured load-displacements for each of the five large footings at TAMU (Briaud amp Gibbens
1994) sand site B-12
Figure B10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU (Briaud amp
Gibbens 1994) footings B-13
Figure B11 Characteristic stress versus square root normalized displacements for TAMU footingsB-15
Figure B12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sandB-16
Figure B13 Summary of sand formation factors with corresponding CPT resistancesB-19
Figure B14 Normalized foundation stresses vs square root of normalized displacement for spread
footings on sand (modified after Mayne et al 2012) B-20
Figure B15 Definitions of capacity from three types of observed foundation load test responses
(modified after Kulhawy 2004) B-21
Figure B16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footingsB-22
B-1
Figure B17 Measured load vs displacement curves for 3 shallow foundations on clay at Baytown Texas
(Stuedlein amp Holtz 2010)B-25
Figure B18 Characteristic stress vs square root of normalized displacement for all three foundations at
Baytown site (Stuedlein amp Holtz 2010) B-26
Figure B19 Normalized foundation stress vs square root of normalized displacements B-27
Figure B20 Applied foundation stress vs CPT-calculated stress for 67 footingsB-28
Figure B21 Applied footing stress vs CPT calculated stress on logarithmic scales B-29
Figure B22 Summary graph and equations of direct CPT method for evaluating the vertical B-30
Figure B23 Foundation soil formation parameter hs versus CPT material index Ic B-32
Figure B24 Bearing capacity ratio qmaxqnet versus CPT material index IcB-33
Figure B25 Influence factor for rectangular foundations from elastic theory solution (Mayne amp
Dasenbrock 2017) B-34
List of Tables
Table B1 Direct CPT methods for bearing capacity of footings on clean sandsB-4
Table B2 Direct CPT methods for bearing capacity of footings on clays B-9
Table B3 Summary of large footings soil conditions and reference sources of database B-17
Table B4 Sources for shallow foundation performace B-34
B-2
B1 DIRECT CPT METHOD FOR SHALLOW FOUNDATIONS
B11 Introduction
The analysis of shallow foundations on soils is classically handled as a two-step and two-part set of
calculations involving (a) bearing capacity commonly reliant on limit plasticity solutions and (b)
settlement or more appropriately termed displacements that are assessed via elastic continuum theory
The utilization of cone penetration tests (CPT) provides the necessary geotechnical data for the site-
specific information on the subsurface conditions at the project site including the geostratigraphy and
evaluation of input parameters This is still a viable approach where the CPT readings are interpreted to
give the soil unit weight (γt) effective friction angle (ϕ) undrained shear strength (su) preconsolidation
stress (σp) and elastic moduli (D and E) for analysis
An alternative approach is the use of CPT results to provide direct assessments of bearing capacity
andor settlements A comparison of these two distinctly different and alternate paths is depicted in
Figure B1
Figure B1 Conventional methods versus direct CPT approach to shallow foundation response
A number of available direct CPT approaches for shallow footings are reviewed within this report
Moreover as the entire load-displacement-capacity response of shallow foundations to loading occurs
as a continuous nonlinear phenomenon a single direct CPT solution is presented for the behavior of
B-3
footings on sands (drained) and clays (undrained) The methods discussed herein are specifically to
address the general case of vertical loading of shallow foundations Additional more complex situations
that consider load eccentricity moments inclined forces sloping ground and other facets may be
handled using well-established procedures that are discussed elsewhere (eg Vesić 1975 Kulhawy et
al 1983)
This report focuses on foundations on granular soils andor soils exhibiting drained behavior since
Federal Highway Administration (FHWA) studies have concluded that less than 1 of shallow
foundations for highway bridges are placed on clay soils (Paikowsky et al 2010) One reason for this is
because soft-firm intact clays require a more complicated process that assesses both a short-term
analysis (undrained) as well as long-term analysis (drained) thus requiring undrained bearing capacity
undrained distortion displacements drained bearing capacity and drained primary consolidation
settlements
B12 Direct CPT Methods for Bearing Capacity of Footings on Sands
Classical bearing capacity solutions are based most often in limit plasticity theory albeit can be
formulated from limit equilibrium cavity expansion and numerical methods Because the limit
plasticity solutions are prevalent and adopt total stress analysis they are usually developed as either
fully drained (Δu = 0) or fully undrained (ΔVV = 0) where Δu equals excess porewater pressure and
ΔVV equals volumetric strain Paikowski et al (2010) provides an extensive review of various solutions
There are a number of direct CPT methods for the evaluation of the bearing capacity of footings and
shallow foundations Table B1 lists a variety of these direct CPT approaches for footings on sands In
the direct approach a one-step process is used to scale the measured cone penetrometer readings (ie
measured cone tip resistance (qc) total cone tip resistance (qt) sleeve friction (fs) andor measured
porewater pressure u2)) to obtain the ultimate bearing stress (qult) of the foundation
Table B1 Direct CPT methods for bearing capacity of footings on clean sands
Method Surface Footing Remarks and Embedment Notes
Meyerhof (1956) qult = qc (B12) middotcw qult = qc (B12)(1+DeB)cw
Note for silty sand reduce by 05
cw = 10 dry or moist sand cw = 05 submerged sand
Meyerhof (1974) qult = qc (B + D)40 with stresses in tsf
Presumably dry sands B = fdn width (feet) and D = depth (feet)
Schmertmann (1978)
NA (applies to embedded footings)
Square qult =055σatm(qcσatm)078
Strip qult =036σatm(qcσatm)078
See Figure A2
Embedment applies De gt 05(1+B) for Blt1m De gt 12 m for B gt1m
B-4
Canadian qult = Rk0middotqc Applied to FS = 3 Geotech Society where Rk0 = 03 where FS = factor of (CFEM 199 2) safety
Tand et al (1995) NA qult = Rkmiddotqc + σvo See Figure A3 where Rk = fctn(De B)
Frank and Magnan (1995)
qult = Rk0middotqc where Rk0 = fctn(sand
consistency) Loose Rk0 = 014
Medium Rk0 = 011 Dense Rk0 = 008
qult = Rk1middotqc + σvo where Rk1 = function (De B L amp
sand consistency) Factor Rk1 = Rk0[1+Rk206+04(BL) (DeB)]
and Rk2 = 035 (loose) 050 (medium) and 085 (dense)
Loose sand qc lt 5 MPa
Medium sand 8 MPa lt qc lt 15MPa
Dense sand qc gt 20 MPa
Eslaamizaad amp Robertson (1996)
qult = KΦmiddotqc See Figure A4
See surface footing equation KΦ = function (BDe shape and density)
Lee amp Salgado (2005)
qbL = βbcmiddotqc(AVG)
where qc averaged over distance B
Not addressed See Figure A5 for factor βbc = fctn(B DR
K0 and sB)
beneath the base
Eslami amp Gholami (2005
2006)
See embedded footing solution
qult = Rk1middotqc where Rk1 = function (ratio DeB
and normalized qcσvo) See Figure A6
Measured qc and qcσvo are geometric means over 2B deep
beneath footing
Robertson amp Cabal (2007)
qult = KΦmiddotqc with KΦ = 016
See surface footing equation
Briaud (2007) qult = KΦmiddotqc with KΦ = 023
Based on full-scale tests at Texas AampM
Mayne amp Illingworth
(2010) qult = 018 qc-Mean
Note qc -Mean is averaged CPT cone resistance over depth of
influence z = 15 B deep
Based on 30 footing load tests on 12
sands
Lehane (2013) qult = 016 qc-AVE Note qc-AVE is averaged CPT cone resistance within z (m) = [B (m)
]07
Based on 47 load tests
Table B1 Continued
Notes B = minimum footing width (or diameter) L = footing length De = embedment depth cw = water
table correction σvo = total overburden stress at bearing elevation and σvo = the effective vertical stress
B-5
Figure B2 Direct bearing capacity relationship for embedded square and strip footings
situated on clean sands (after Schmertmann 1978)
Figure B3 Direct CPT bearing factor Rk as function of footing width B and embedment depth
from finite element analyses (Tand et al 1995)
B-6
Figure B4 Direct CPT bearing factor Rk as function of footing shape size-to-embedment ratio
and sand consistency (Eslaamizaad amp Robertson 1996)
B-7
Figure B5 Direct CPT factors for bearing capacity of sands from FEM analysis by Lee amp
Salgado (2005) in terms of footing size relative density and base settlement
Figure B6 Direct CPT bearing factor Rk as function of footing embedment to size ratio and
normalized cone tip resistance (after Eslami amp Gholami 2005)
B13 Direct CPT Methods for Bearing Capacity of Footings on Clays
For direct CPT evaluations of foundation bearing capacity in clays Table B2 summarizes some of the
well-documented methods that are available Generally these assume that the loading is fast enough
and the permeability of the clay is also sufficiently low such that an undrained (ie constant volume)
condition is maintained This is fine for short term loading of foundations particularly for soft to firm
intact clays However the undrained case is not permanent Given sufficient time excess porewater
pressures in the clay will eventually dissipate and hydrostatic conditions will return to equilibrium
During this dissipation phase primary consolidation will occur that results in drained settlements Also
a drained bearing capacity condition will prevail As the CPT is advanced at a constant rate of 20 mms
the recorded readings normally constitute undrained behavior from a direct measurement viewpoint
Thus direct CPT methods are normally focused at an undrained foundation response Nevertheless the
drained footing case can be addressed using CPT results via use of conventional analysis approach and
effective stress limit plasticity solutions that require piezocone penetrometers with porewater pressure
measurements
B-8
B
D)
L
B4060(3501320kc
Many of the earlier solutions were based on data from mechanical penetrometers andor older electric
cones that measure the uncorrected tip resistance (qc) It is now well-recognized that this measurement
must be updated to the corrected cone resistance (qt) because of porewater pressure effects that act on
the mechanics of the load cell and geometry of the particular penetrometer design The results are
particularly significant in clays silts and mixed soils that are soft to firm to stiff where excess porewater
pressures are recorded
Table B2 Direct CPT methods for bearing capacity of footings on clays
Method Footing Comments NotesRemarks
Meyerhof (1974) qult = αbcmiddotqc
where 025 le αbc le 050
Applicable to saturated insensitive clays under short-term loading
Cone tip resistance
measured by
mechanical CPT
Trofimenkov (1974)
Approximation
qult asymp (qc33)09
(Assuming FS = 3)
Strip footings on clays and sandy clays 06m le B le 15m 1m le zemb le 25m
Mechanical CPT
with qc and qult in
kgcm2
Schmertmann (1978)
qult = function (qc and foundation shape)
Relationship shown in Figure A7 for square and strip footings
Cone tip resistance
measured by
mechanical CPT
Tand et al (1986)
qult = Rkmiddot(qc - σvo) + σvo
Factor Rk obtained from Figure A8 accounts for surface and deep footings Separate Rk factor for intact and jointed clays
= (qc1middotqc2)05 qc
where qc1 is geometric mean from bearing elevation to 05B deeper and qc2 is geometric mean from 05B to 15B beneath foundation base
Mix of data from
mechanical CPT qc
and electric CPT qc
LCPC Method
(Frank amp Magnan 1995)
Footing a surface
qult = kc middotqc + σvo
where kc = 032
Embedment case
Note Methods above generally consider undrained bearing capacity
B-9
Figure B7 Direct relationship between ultimate bearing stress in clays and measured cone tip
resistance (Schmertmann 1978)
B-10
Figure B8 Direct relationship between ultimate foundation bearing stress in clays and net
cone tip resistance (after Tand et al 1986)
B14 Direct CPT Assessments of Settlement and Displacements of Shallow Foundations
Methods for evaluating the magnitude of foundation displacements (s) can be found using
elastic theory subgrade reaction methods spring models and numerical simulations The
classical approach to footing settlement calculations on sands is to utilize elastic theory
solutions (eg Poulos amp Davis 1974) which take on the form
119902∙119861∙119868∙(1minus1205842) 119904 = B1
119864119904
where q = applied footing stress and I = displacement influence factor from elasticity theory
The value of I depends upon foundation geometry footing rigidity embedment depth variation of soil
modulus with depth compressible layer thickness and other factors as discussed by Mayne amp Poulos
(1999) For instance for a flexible circular footing of diameter B resting on an infinitely deep soil
formation having a constant modulus Es with depth the influence factor I is equal to 1 for the center
point The results of in-situ field tests such as pressuremeter tests (PMT) standard penetration tests
(SPT) cone penetration tests (CPT) flat dilatometer tests (DMT) and other methods can be used to
ascertain the input value of soil modulus Es
Alternatively direct CPT methods have been investigated to provide a one-step assessment of
foundation displacements Selected methods for direct CPT evaluation of shallow foundation
settlements on granular soils is include DeBeers amp Martens (1957) Meyerhof (1965) DeBeer (1965)
Thomas (1968) Schmertmann (1970) Meyerhof (1974) Berardi amp Jamiolkowski amp Lancellotta (1991)
Robertson (1991) Lutenegger amp DeGroot (1995) Lehane amp Dougherty amp Schneider (2008) Gavin amp
Adekunte amp OrsquoKelly (2009) Mayne amp Illingworth (2010) and Uzielli amp Mayne (2012)
B141 Large Scale Footing Load Tests
The measured load-displacement response of shallow foundations is conducted in the field using either
stepped loading procedures or continuous rate of displacement methods Results from the well-
documented test program at the Texas AampM University (TAMU) site (Briaud amp Gibbens 1999 Briaud
2007) for five spread footings on sand are shown in Figure B9 Each footing clearly shows a nonlinear
B-11
behavior to loading over the range of testing This site is one of the national geotechnical
experimentation sites (NGES) funded by the National Science Foundation (NSF) FHWA and American
Society of Civil Engineering (ASCE)
Figure B9 Measured load-displacements for each of the five large footings at TAMU (Briaud
amp Gibbens 1994) sand site
B142 Generalized Direct CPT Method for Footing Response on Soils
The use of applied foundation stress versus normalized displacement curves (q vs sB) is generalized to
footings on sands silts intact clays and fissured clays A square root plotting of the normalized
displacements (sB) permits a single parameter characterization for each specific soil type that in turn
correlates with the net cone tip resistance The generalization is based on a statistical review of data
from large foundations (B gt 05m) involving 70 full-scale load tests with available CPT data For fine-
grained soils the data are from electronic piezocone tests where the proper correction from qc to qt has
been obtained to allow the best possible results The methodology permits an evaluation of the load-
displacement-capacity response of shallow square footings based on CPTs performed in the four types
B-12
of ground conditions For the TAMU footings the characteristic stress-normalized displacement
response is shown in Figure B10 where all five foundations can be represented by a single curve
Figure B10 Characteristic nonlinear stress-normalized displacement curve for the five TAMU
(Briaud amp Gibbens 1994) footings
B143 Characteristic Stress-Displacement Curves
Fellenius amp Altaee (1994) recommended the concept of a characteristic stress (q) vs normalized
displacement (sB) curve for foundation response on a given soil formation later supported by Decourt
(1999) Briaud amp Gibbens (1999) Lutenegger amp Adams (2003) and Briaud (2007) In this approach the
measured load (Q) vs displacement from individual footings of various sizes that rest on the same soil
conditions collapse to a unified relationship given in Equation B2
119939119943 119956 119954119938119953119953119949119946119942119941 = 119938119943 (119913
) B2
B-13
where af and bf are empirical fitting coefficients (Decourt 1999 Uzielli amp Mayne 2011 2012)
In a review of measured load tests data from large spread footings situated on different sands it has
been suggested that Equation B2 can be reduced to a single parameter expression by use of square root
plotting where the exponent bf is equal to 05 (Mayne amp Illingworth 2010 Mayne et al 2012)
119956 119954119938119953119953119949119946119942119941 = 119955119956radic
119913 B3
where rs = fitted soil parameter from best fit line regression In the case of sands the coefficient rs
represents the supporting soil conditions including particle size relative density and fines content as
well as geologic origin aging and overconsolidation effects
The results from the five TAMU footings are presented in this format in Figure B11 that shows the
formation factor rs = 486 MPa for this sand deposit This single coefficient can be used to express the
nonlinear load versus settlement of any size footing on this sand site Moreover one criterion from
Eurocode for bearing capacity that is used by the European community identifies that stress (or load)
which corresponds to (sB) = 10 That is when the displacement equals ten percent of the footing
width then the capacity has been reached Therefore adopting this criterion for footings on sands
the ultimate bearing stress (qult) for footings on sand can be taken when (sB) equals 010 or when
square root of (sB) equals 0316 In that case this gives qult equal to 0316 middot rs For the TAMU site this
gives qult equal to 0316 times (486) which equals 153 MPa as illustrated by Figure B12
B-14
Figure B11 Characteristic stress versus square root normalized displacements for TAMU
footings
B-15
Figure B12 Eurocode or LCPC criterion applied to determine bearing capacity of TAMU sand
B15 Footing Database
A database of spread footings and large plates was assembled which considers only full-size shallow
foundations (05 le B le 6 m) that rest on sands silts and clays Sites for the foundations were also subjected to CPT The inclusion solely of large footings are significant because results from small-scale
model footings exhibit scale effects In fact experimental centrifuge work and numerical simulation
studies have shown that bearing capacity factors are size-dependent especially for factor Nγ decreasing
with footing width B (Kimura et al 1985 Cerato amp Lutenegger 2007 Mase amp Hashiguchi 2009) A
direct approach based on full-size footings provides a more reliable means for foundation evaluation
The compiled database is given in Table B3 with a total of 70 large footings and plates These include a
listing of 34 foundations on 13 sands 11 footings on 4 silt deposits 13 footings or large plate load tests
on 6 intact clays and 12 foundations on 5 fissured clays Most of the footings were square or nearly
square (80) while the remaining were circular The largest foundation was a mat 5 m by 14 m in plan
loaded by stacking Kentledge concrete blocks to cause a bearing failure in the underlying soft clay For
sands the largest footing consisted of the 6-m square concrete pad In terms of equivalent square
footings mean footing sizes were 155 m (sands) 114 m (silts) and 126 m (clays)
B-16
Additional details on the individual load tests and site conditions are given elsewhere (Mayne 2009
Mayne amp Illingworth 2010 Uzielli amp Mayne 2011 2012 Mayne et al 2012 Mayne amp Woeller 2014)
Table B3 Summary of large footings soil conditions and reference sources of database
Sand Site Location Soil Conditions Footings Numbers
Shapes and Sizes
ReferencesSource
College Station
Texas Pleistocene sand 5 Square 1 15 25 33 m Briaud amp Gibbens (1999)
Kolbyttemon Sweden Glaciofluvial sand 4 Rect B = 06 12 17 24 m
Bergdahl et al (1985 1986)
Fittija Sweden Glaciofluvial sand 3 Rect B = 06 17 24 m Bergdahl et al (1984 1985)
Alvin West Texas Alluvial sand 2 Circular D = 235 m Tand et al (1994)
Alvin East Texas Alluvial sand 2 Circular D = 22 m Tand et al (1994)
Perth Australia Silceous dune sand 4 Square B = 05 and 10 m
Lehane (2008)
Grabo T1C Sweden Compacted sand fill
1 Square B = 046 m Long (1993)
Grabo T2C Sweden Compacted sand fill
1 Square B = 063 m Long (1993)
Grabo T3C Sweden Compacted sand fill
1 Square B = 080 m Long (1993)
Labenne France Dune sand 6 Square B = 07 and 10 m
Amar et al (1998)
Green Cove Florida Brown silty sand 1 Circular D = 182 m Anderson et al (2006)
Durbin South Africa
White fine sand 1 Square B = 609 m Kantley (1965)
Porto Portugal Residual clayey sands
1 Circular D = 12 m and 1 plate
Viana da Fonseca (2003)
Jossigny France Soft clayey silt 2 Square B = 1 m Amar et al (1998)
Tornhill Sweden Glacial Baltic till 3 Square B = 05 1 2 m Larsson (2001)
Vagverkel Sweden Stiff medium silt 3 Square B = 05 1 2 m Larsson (1997)
Vattahammar Sweden Brn-Gry layered silt 3 Square B = 05 1 2 m Larsson (1997)
B-17
Table B3 Continued
Clay Site Location Soil Conditions Footings Numbers Shapes and Sizes
ReferencesSource
Baytown Texas Fissured Beaumont 2 plates with d = 076 m Stuedlein amp Holtz (2008)
Belfast Ireland Soft clay silt ldquosleechrdquo
1 Square Pad B = 20 m Lehane (Geot Engr 2003)
Bothkennar Scotland Soft silty clay 2 Square Pads B = 22 24 m
Jardine et al (1995)
Bangkok Thailand Soft to stiff clay 4 Square 067 075 090 10 m
Brand et al (1972)
Haga Norway Stiff OC clay 2 Square Footings B = 10 m
Andersen amp Stenhammar (1982)
Rio Grande Brazil Sandy residual clay 3 Square 04 07 and 10 m
Consoli et al (1998)
Shellhaven England Soft estuarine clay Rectangular 5 m by 14 m Schnaid et al (1993)
Texas City A Texas Coast Fissured Beaumont 3 circular plates d = 058 m
Tand et al (1986)
Texas City B1 Texas Coast Fissured Beaumont 2 circular plates d = 058 m
Tand et al (1986)
Texas City B2 Texas Coast Fissured Beaumont 1 circular plate d = 058 m
Tand et al (1986)
Alvin Texas Texas Coast Fissured Beaumont 3 circular plates d = 058 m
Tand et al (1986)
For the series of footings on sands Figure B13 shows the summary of measure formation factors (rs) for
the 34 foundation load tests Also indicated on the graph are the corresponding mean qc values of the
sands averaged over 15middotB beneath the bearing elevations
Note also in sands that the qt and net resistance (qnet) are all very close to the measured qc since (a)
porewater pressure corrections for the a-net value are negligible in clean sands and (b) overburden
stresses are typically only 1 to 3 of the magnitude of qc This is especially true of shallow depths
Therefore for clean sands qnet is approximately qt which is approximately qc
B-18
Figure B13 Summary of sand formation factors with corresponding CPT resistances
As suggested by Briaud (2007) various in-situ measurements can be used to normalize the results from
cone penetration tests in this case Herein the qc from the CPT soundings in the sands are used to
normalize the footing stress axis as presented in Figure B14 It is evident that the results of the load-
displacement response of all 34 footings on 13 different sands can be captured by the characteristic
stress versus normalized displacement with the inclusion of the site-specific cone tip resistance In this
case the mean trend for all footings and sand sites indicates a simple expression shown in Equation B4
119954119943119952119952119957119946119951119944 119956 Clean quartz-silica sands = 120782 120787120790120787radic B4
119954119940 119913
B-19
Figure B14 Normalized foundation stresses vs square root of normalized displacement for
spread footings on sand (modified after Mayne et al 2012)
The results of measured force-displacement curves (Q vs s) from footing load tests are nonlinear and
thus the definition of capacity must be addressed Kulhawy (2004) identified three main curve types
that occur during load testing depicted in Figure B15 The most common response (Type A) shows no
clear peak and is observed in sands and silts as well as slow loading of insensitive clays and fissured
clays In this case ldquocapacityrdquo can be defined by the Eurocode corresponding to the load (or stress) when
sB=10 (Amar et al 1998) For foundations situated on many clays a plateau capacity is reached as
depicted as Type B response In rare cases where sensitive clays or structured soils are encountered a
Type C relationship occurs whereby the load-displacement curve reaches a peak followed by strain
softening In the one case observed in this study for Type C loading (Haga clay) the corresponding
ldquopeak capacityrdquo was reached at a pseudo-strain of sB = 4 as shown in Figure B16 followed by some
strain softening In the cases of soft clays at Belfast and Bothkennar a plunging type (plateau failure)
was achieved at around sB=7
B-20
Figure B15 Definitions of capacity from three types of observed foundation load test
responses (modified after Kulhawy 2004)
B-21
Figure B16 Summary of applied footing stresses normalized to CPT net cone resistance vs pseudo-
strains (sB) for all footings
In the case of fine-grained soils that develop excess porewater pressure during cone penetration the
use of the net total cone resistance (qtnet) which equals the cone tip resistance minus the total stress
must be considered (Lunne et al 1997) As noted earlier the correction of CPT data in sands is minimal
because of the low value of penetration porewater pressures and the fact that total overburden stress is
small relative to the cone resistance especially for shallow soundings Thus the use of qtnet may be
substituted for qc into the trend of Figure B1
B16 Undrained and Drained Capacity
The footings on sands typically show drained response with no excess porewater pressures developed
during loading For the foundations situated on silts analyses by Larsson (1997 2001) concluded that
these too show drained conditions For shallow foundations on clays the entire range of drainage cases
are possible ranging from undrained to partially-drained to fully-drained depending upon the applied
rate of loading relative to the permeability of the geomaterial If sufficient data were available for each
of the clay sites then the recent evaluation of drainage condition could be assessed using normalized
velocity criteria (Chung et al 2006) However this was not possible with the current data set Instead
B-22
the test loadings of footings on clays have essentially been assumed to primarily occur under undrained
conditions following recommendations by Tand et al (1986) For this case a limiting bearing stress can
be calculated from bearing capacity analyses and related directly to the CPT readings
From classical bearing capacity calculations using limit plasticity solutions the ultimate foundation stress
is shown in Equation B5
= 119925119940 ∙ 119956119958 B5 119954119958119949119957
where Nc equals the bearing factor for constant volume (514 for strip and 614 for square or circular
plan) and su equals the undrained shear strength of the clay For the case of soft to firm clays of low to
medium sensitivity the operational strength can be related to the preconsolidation stress (Mesri 1975
Jamiolkowski et al 1985) as shown in Equation B6
prime 119956119958 = 120782 120784120784120648119953 B6
Finally the effective preconsolidation stress has been linked directly to the net cone resistance (Chen amp
Mayne 1996 Demers amp Leroueil 2002) as shown in Equation B7
prime 120648119953 = 120782 120785120785(119954119957 minus 120648119959119952) B7
Combining Equations (B5) (B6) and (B7) results in direct expressions for undrained bearing capacity on
intact clays from CPT results as shown in Equations B8 and B8-2
Strip footing 119954119958119949119957 = 120782 120785120789120785(119954119957 minus 120648119959119952) B8
B-23
Squarecircle 119902119906119897119905 = 0445(119902119905 minus 120590119907119900) B8-2
Equation (B8-2) is in excellent agreement with Tand et al (1986) database study for the case of intact
clays and footings with no embedment Their study also provided a lower relationship for foundations
situated on jointed clays specifically recommending that the bearing stress not exceed 030 qtnet
Review of the available data in Figure 3 shows this to be rather conservative and a more realistic value
may be taken at 040 qtnet for fissured clays
B161 Normalized Undrained Footing Response
The characteristic stress vs normalized displacement curves for footings on clays exhibiting undrained
conditions can be verified using a relatively recent case study on Beaumont clay by Stuedlein amp Holtz
(2010) The Baytown Texas site was investigated using an exploration program of soil borings
piezocone soundings and laboratory testing The field testing included a large square footing (B = 276
m) and two small circular plates (D = 076 m) The measured load-displacement responses for all three
foundations are presented in Figure B17 Of course the large footing carried considerably higher loads
than the two small plates on the same soil deposit Yet using the aforementioned characteristic stress
(q) vs square root of normalized displacement (sB) essentially gave a unique relationship for all 3
foundations as presented in Figure B18 In this case utilizing the form of Equation B2 the
characteristic coefficient rs is 177 MPa for this deltaic clay formation
B-24
Figure B17 Measured load vs displacement curves for 3 shallow foundations on clay at
Baytown Texas (Stuedlein amp Holtz 2010)
B-25
Figure B18 Characteristic stress vs square root of normalized displacement for all three
foundations at Baytown site (Stuedlein amp Holtz 2010)
B17 General Direct CPT Method
In a manner analogous to the normalization of applied foundation stresses shown in Figure B14 for
footings on sands a generalized direct CPT method for shallow foundations on different soils can be
made in Equation B9
120782120787 119956 119954 = 119945119956 ∙ 119954119957119951119942119957 ∙ (119913
) lt 119954119940119938119953119938119940119946119957119962 B9
where qcapacity is equal to the foundation bearing capacity of the ground and hs is equal to the empirical
fitting term that depends on soil type Specifically hs equals 058 for sands 112 for silts 147 for
fissured fine-grained soils and 270 for clays A summary of the normalized stress versus the normalized
displacement curves for these four soil types is presented in Figure B19 where the data fitting to obtain
the hs parameter on clays are limited to (sB) lt 004 corresponding to undrained loading The data on
B-26
silts and sands are considered fully drained whereas the fissured clay subset may be partially drained to
undrained The statistical measures for obtaining the fitted parameter hs are quite good as shown in
Figure B20 with the coefficient of determinations (r2) values of 0947 for sands 088 for silts 0935 for
fissured and 0925 for intact clays Additional statistical evaluations on the database are provided by
Uzielli amp Mayne (2011)
Figure B19 Normalized foundation stress vs square root of normalized displacements
B-27
Figure B20 Applied foundation stress vs CPT-calculated stress for 67 footings
The same data are shown on Figure B21 in a log-log plot to emphasize the early parts of the load-
displacement responses of these footings The best fit line from regression analyses shows excellent
statistical correlations As detailed earlier the undrained bearing capacity of square or circular footings
on clays can be taken as approximately 040 times qtnet For silts and sands that experience drained
loading with no excess porewater pressures being developed and no clear peak value for capacity the
(sB) =10 criterion can be used In this case Equation 7 can be used directly to obtain stress q equal to
qcapacity when (sB)05 is 0316
An alternate approach is presented in Figure B22 summarizing the direct CPT method for estimating the
load-displacement-capacity of shallow square footings on four soil types The maximum values of soil
bearing capacity (qmax) can be capped at stresses corresponding to a percentage of the average
measured qtnet determined over the depth interval from the foundation bearing elevation to 15middotB
below the foundation In such an approach the foundation bearing capacity can be taken as
(qcapacityqtnet) of 02 for sands 035 for silts 040 for fissured and 045 for intact clays Using the assigned
values of the hs parameter the corresponding cut-off for the general stress-normalized displacement
relationship given by Equation 7 can be established specifically giving corresponding values of the
B-28
normalized displacements which can also be considered as pseudo-strains equal to (sB) max of 4 for
clays 7 for fissured clays 10 for silts and 12 for sands
Figure B21 Applied footing stress vs CPT calculated stress on logarithmic scales
B171 CPT Soil Material Index Ic
The soil behavioral type can be assessed indirectly by CPT using a material index (Ic) as defined by
Robertson (2009a b) shown in Equation B10
119920119940 = radic[120785 120786120789 minus 119949119952119944119928119957119951]120784 + [120783 120784120784 + 119949119952119944119917119955]120784 B10
B-29
Figure B22 Summary graph and equations of direct CPT method for evaluating the vertical
where Qtn equals the stress-normalized cone tip resistance and Fr equals the normalized sleeve friction
calculated from the cone penetrometer readings as shown in Equations B11 and B12 respectably
(119954119957minus120648119959119952)120648119938119957119950 = 119951 B11 119928119957119951 prime (120648119959119952120648119938119957119950)
119943119956 119917119955() = 120783120782120782 ∙ B12 (119954119957minus120648119959119952)
where σatm equals a reference stress equal to atmospheric pressure (σatm = 1 atm asymp 1 bar asymp 100 kPa) In
the initial evaluation the exponent n is set to n = 1 to find the soil behavioral type (SBT) based on a 9-
zonal chart as shown in Figure 23 The value of exponent n varies with soil type ranging from n = 1 in
intact clays and decreasing with increasing grain size to around approximately 075 in silts and
B-30
approximately 05 plusmn 02 in clean sands The appropriate value of exponent n is found by iteration until
conversion using the relationship (Robertson 2009b) shown in Equation B13
prime 120648119959119952 119951 = 120782 120785120790120783 ∙ 119920119940 + 120782 120782120787 ( ) minus 120782 120783120787 le 120783 120782 B13 120648119938119957119950
As the distinction that footings on intact clays subjected to fast loading are behaving primarily under
undrained conditions (ie constant volume) Robertson (2009a) suggested that this occurs when Ic gt 27
while in contrast Ic lt 25 corresponds more or less to drained loading (ie no excess porewater
pressures)
The CPT material index can be used to identify soil type and the corresponding characteristic hs value for
estimating footing load-displacement response In a number of the case studies investigated the results
of the sleeve friction readings were also available to allow the calculation of Ic with depth at these sites
Figure B23 shows the tentative relationship between the values of the hs parameter plotted versus the
corresponding CPT material index with an approximate trend given by Equation B14
120784120785 119945119956 = 120784 120790 minus 120783120787 B14
119920119940 120783+( ) 120784120786
B-31
Figure B23 Foundation soil formation parameter hs versus CPT material index Ic
Soil behavioral type from the Ic ranges given by Robertson (2009b) and are shown in Figure B23 with an
overall general agreement with the known soil classifications In summary the CPT can provide an
evaluation of the load-displacement-capacity response directly via Equation B15 which may be of
interest in mixed soil types such as silty sands sandy silts and the like
120784120785 Footing stress 119954 = 119954119951119942119957 ∙ radic(119956frasl119913) ∙ [120784 120790 minus
)120783120787] B15
120783+(119920119940frasl120784120786
As discussed earlier foundation bearing capacity may be taken by a limiting value of pseudo-strain (sB)
max or by qmax as specified as a percentage of qtnet This definition of bearing capacity can be seen in
comparison to soil type as seen in Figure B24
B-32
Figure B24 Bearing capacity ratio qmaxqnet versus CPT material index Ic
B172 Rectangular Foundations
This same elastic solution from Giroud (1968) for a CPT method for square and circular foundations were
utilized for rectangular foundations The study covered rectangular distortions (AB) ranging from 1
(square) to very long foundations with AB equal to 20 where A represents the foundation length and B
represents the foundation width (Mayne amp Dasenbrock 2017) The influence factor for rectangular
shaped foundations is given in Figure B25 and the expression in Equation B16
= (119912119913)120782120785120786120787 119920119912119913 B16
B-33
Figure B25 Influence factor for rectangular foundations from elastic theory solution (Mayne
amp Dasenbrock 2017)
Including 32 full scale load tests on sands results from an additional 98 shallow foundations have been
reported in previous studies The foundations were primarily square or circular and among these include
large spread footings with measured settlements and bearing capacity Table B4 shows a summary of
the number of footings and footing sizes used in the study The range of length to width ratios varied
from 1 lt (AB) lt 23 with an average value of (AB) equal to 238 The embedment depth to footing
width ranged from 0 lt DeB lt 222 and averaged 046
Table B4 Sources for shallow foundation performace
Source of Data No of
Footings
Mean B
(m)
Max B
(m)
Min B
(m)
Mean A
(m)
Max A
(m)
Min A
(m)
Mayne et al (2012) 32 149 609 046 149 609 046
Lehane (2011) 8 041 027 060 041 060 027
Gifford et al (1987) 17 846 1588 527 1591 3596 701
Jeyapalan amp Boehm
(1986)
13 1022 2892 400 1671 3049 640
B-34
Schmertmann
(1970)
31 945 5608 061 1288 8668 061
Papadopoulos
(1992)
29 870 3600 100 1300 7290 100
Total 130 671 5608 027 1012 8668 027
The solution for shallow foundation settlements given back in Equation B1 combined with
the expression in Equation B16 takes on the form
119954∙119913∙119920119912119913∙(120783minus120642120784) 119956 = B17
119916119956
Combining the direct CPT equation developed from the 32 full scale load tests (Equation B15)
together with the influence factor for rectangular shaped foundations becomes
minus120782120785120786120787 120784120785 119912 Footing stress 119954 = 119954119951119942119957 ∙ radic(119956frasl119913) ∙ [120784 120790 minus
)120783120787] ∙ [ ] B18
120783+(119920119940frasl120784120786 119913
B18 Conclusions
A direct CPT method for square rectangular and circular shallow footings is developed using a database
of 166 full-scale field load tests where the minimum footing size of an equivalent square width of 05 m
is established to avoid scaling issues The use of load vs displacement is generalized by characteristic
stress vs normalized displacement (sB) and further simplified by a square root plotting technique that
captures the ground response in a single parameter that is related to the qtnet Data are grouped
according to four main soil categories sands silts fissured clays and intact clays It is generally believed
that the footings on sands and silts exhibit fully drained behavior while in intact clays an undrained
response occurs under conditions of constant volume
B-35
The summary equation for evaluating the vertical stress-displacement-capacity of square rectangular
and circular footings is given by
minus120782120785120786120787 119956 119912 119954119943119952119952119957119946119951119944 = 119945119956 ∙ 119954119957119951119942119957 ∙ radic ∙ ( ) lt 119954119950119938119961 B19
119913 119913
where the parameter hs also relates directly to Ic The foundation capacity depends upon the mode of
failure as well as drainage conditions and can be taken as a fraction of qtnet where qmaxqtnet is 020 in
sands 035 in silts 040 in fissured clays and 045 in intact clays as shown in Figure 24 Alternatively the
capacity can be defined using a limiting pseudo-strain given by (sB) max of 4 in clays 7 in fissured
10 in silts and 12 in sands
B-36
APPENDIX C
List of Figures
Figure C1 Components of Axial Pile Capacity C-3
Figure C2 Selected alpha relationships as function of the undrained shear strengthC-5
Figure C3 Pile side friction coefficient β in terms of effective friction angle and overconsolidation ratio
C-11
Figure C4 Concept of Direct versus Rational Method for CPT evaluation of axial pile capacityC-14
Figure C5 Soil behavior type using CPT via original UniConeC-19
Figure C6 Soil behavior type using CPT via Modified UniCone C-19
Figure C7 Nine-zone soil behavioral soil type using normalized piezocone parameters C-21
Figure C8 Summary of Modified UniCone Method for direct CPT assessment of axial pile capacity C-22
Figure C9 Elastic continuum solution for axial pile response under compression and tension loadingC-24
List of Tables
Table C1 Alpha Methods for Axial Pile Side Friction where fp = αmiddotsu C-5
Table C2 Beta Methods for Axial Pile Side Friction where fp = βmiddotσvoC-9
Table C3 Selection of Direct CPT Methods for Axial Pile CapacityC-15
C-1
C1 EVALUATING DEEP FOUNDATION RESPONSE FROM CONE
PENETRATION TESTS
C11 Introduction
The axial response of deep foundations includes the load-displacement-capacity and axial load transfer
when driven pilings and drilled shafts are loaded in compression and uplift The use of cone penetration
testing (CPT) especially seismic piezocone tests (SCPTu) are advantageous since they provide
information on the subsurface soils and their geomaterial properties The SCPTu provides at least four
measurements on soil behavior with depth including (a) cone tip resistance qt (b) sleeve friction fs (c)
penetration porewater pressure u2 and (d) shear wave velocity Vs This offers the opportunity to
determine the geostratigraphy unit weight effective overburden stress shear strength stress state
and stiffness of the ground from a single exploratory sounding thus values in economic expediency
and reliability The axial pile capacity can be calculated based on static equilibrium of forces acting along
the sides and base of the pile foundation The displacement of the pile can be ascertained using
elasticity theory from closed-form solutions boundary elements andor finite element analyses Load
transfer occurs along the length of the pile and only a portion of axial forces are transmitted to the base
or toe or tip of the pile These too can be assessed using elasticity solutions
An alternate approach to pile capacity involves the utilization of direct CPT methods available
since the 1970s but in the past 10+ years a number of new and statistically reliable algorithms
have been developed which can be implemented for highway design and construction
C111 Axial Pile Capacity
The axial compression capacity of a single pile foundation is composed of a shaft or side component and
end-bearing component at the base as depicted in Figure 1 For a circular pile the side capacity (Qs) is
determined from the unit side friction (fp) acting along the surface area of the shaft which is As = πmiddotdmiddotL
where d = pile diameter and L = length embedded below grade
If the magnitude of side friction is uniform and constant with depth the side capacity is simply
C-2
119928119956 = 119943119953 middot 119912119956 C1
Moreover however many piles extend through multiple layers and a summation of unit side
frictions acting on various pile segments must be tabulated over the length of the pile as
suggested by Figure C1
Figure C1 Components of Axial Pile Capacity
The unit-end bearing resistance (qb) acts over the base of the pile tip where the area of a circular pile is
given by Ab = πmiddotd24 For piles in compression loading the base capacity is determined from Equation
C2
119876119887 = 119902119887 middot 119860119887 C2
and for piles in tension (or uplift) it is normally taken that Qb = 0
C-3
C112 Pile Unit Side Friction
Several different approaches can be adopted for evaluating the unit pile side friction (fp) prior to full-
scale load testing and construction (Poulos amp Davis 1980 ONeill 2001) The most common types
including the alpha and beta methods as summarized in Tables 1 and 2 respectively In the alpha
method an empirical coefficient (α) is applied to the undrained shear strength (su) of clay soils to
obtain
119891119901 = 120572 middot 119904119906 C3
An illustrative example of alpha curves is shown in Figure C2 A difficulty with the alpha
method is the evolution of many variants and changes to the expressions for its estimation It is
also restricted in its use for specific types of piles in clays and fine-grained soils
C-4
Figure C2 Selected alpha relationships as function of the undrained shear strength
Table C1 Alpha Methods for Axial Pile Side Friction where fp = αmiddotsu
C-5
Reference
Source
Alpha Equation Remarks
Tomlinson
(1957)
2716801102
uu ssAll pile types (steel
concrete timber)
Note su in ksf
Tomlinson
(1957)
2616201102
uu ssConcrete piles
Note su in ksf
McClelland
(1974)
Kerisel α = 07 - 031middotln(su)
Peck α = 10 - 02middotsu
Woodward α = 091middot(su)091
Driven piles in clay
Note su in ksf
Semple
(1980) )1(511
)1(1
OCR
OCR Summary from 9 series of
pile load tests
American
Petroleum
Institute (API
1981)
For suσvo le 035 α = 10
For suσvo gt 080 α = 05
Otherwise α = 1 + 1111middot (035 - suσvo)
Driven steel pipe piles
Tomlinson
(1986)
1 α = 55 (su)-091
2 α = 785 (su)-102
3 α = 605 (su)-100
where 75 lt su (kPa) lt 210
1 Driven piles
2 Bored piles
3 Driven piles in till with L
lt 10middotd
API (1987) 1 α = 10 for su lt 25 kPa Driven piles in clay other
than Gulf of Mexico
C-6
2 α = 125 - 001middotsu for 25 lt su lt 75 kPa
3 α = 050 for su gt 75 kPa
Kulhawy amp
Jackson
(1989)
α = 021+026middot(σatm su) le 1 Analyses of 106 drilled
shaft foundations
Table C1 Continued
API (1989) 05 For suσ vo le 1 α =
su radic frasl prime σvo
minus025 su For suσ vo gt 1 α = 05 ∙ ( frasl prime ) σvo
Driven steel pipe piles
Kulhawy amp
Jackson
(1989)
OCR
OCR
sin50
tan)sin1( sin
Λ = (1-CsCc) asymp 080 for
insensitive clays
Nowacki et al
(1996)
minus05 su For suσvo le 07 α = 05 ∙ ( frasl prime ) σvo
minus02 su For suσvo gt 07 α = 055 ∙ ( frasl prime ) σvo
Driven pile foundations
Kolk and van
der Velde
(1996)
α =09middot FL middot (suσvo)-03 lt 10
FL = [ (L - z)d]-02 = length term
L = pile length
Note su obtained from
UU lab tests
Applicable to driven piles
in clays
C-7
Knappett amp α = 1 for su le 30 kPa Non-displacement piles in
Craig (2012) fine-grained soils α = 116 - (su185) for 30 kPa le su le 150 kPa
α = 035 for su gt 150 kPa
Knappett amp
Craig (2012) α = 055 ∙ 02 40
( ) (LfraslD)
∙ su ( frasl prime σvo
minus03
) Displacement piles in fine-
grained soils
z = depth at point considered
d = outside diameter of pile
Karlsrud et al
(2005 NGI
Method)
α = function (suσvo and plasticity index) Driven pilings
Fleming et al
(2009) 05 minus05 su su For su σvo le 1 120572 = ( frasl prime ) ∙ ( frasl prime )
σvo NC σvo
05 minus025 su su For su σvo gt 1 α = ( frasl prime ) ∙ ( frasl prime ) σvo NC σvo
For driven piles in clay
where (suσvo)NC is the
normalized shear strength
for normally-consolidated
clay
Brown et al
(2010)
α = 0 for 0 lt z le 5 feet
α = 055 for z gt 5 feet and (suσatm) le 15
α = 055-01middot(suσatm -15) for 15le (suσatm) le 25
Drilled Shafts and Bored
Piles
Table C1 Continued
C-8
OCRK )sin1(0
Karlsrud
(2012)
α = function (su σvo and plasticity index) Driven pilings
Note as reported by Karlsrud (2012)
In the beta method the coefficient β is applied to the effective overburden stress (σvo) at the point of
concern along the pile length
prime = 120573 C4 119891119901 middot 120590119907119900
Table C2 Beta Methods for Axial Pile Side Friction where fp = βmiddotσvo
Reference Source Beta Equation Remarks
Burland (1973) β = (1-sinϕ) middot tan ϕ Piles in NC clays
Meyerhof (1976) β = Kmiddottanδ where K = K0 = for NC clays
K = 15middotK0 for OC clays
Poulos amp Davis
(1980)
β = (1-sinϕ) middot tan ϕ middot OCR05 Piles in OC stiff clays
Table C2 Continued
C-9
Kulhawy amp Jackson
(1989)
β = (1-sinϕ) middot tan ϕ middot OCRsinϕ Piles in quartz sands and insensitive
clays
ONeill (2001) β = (1-sinϕ) middot tan θ middot OCRsinϕ Drilled shafts where θ = (δϕ)middotϕ and
δ = interface friction between pile and soil
Fleming et al
(2009)
β = Kmiddottanδ K = lateral stress coefficient and δ =
friction angle between soil and pile
material
Karlsrud (2012) β = fctn (OCR and PI) PI = plasticity index of the clay ()
Mayne and Niazi
(2017)
β = CM middot CK middot K0 middot tanϕ
where K0 = (1-sinϕ) middot OCRsinϕ for soils
that are virgin loaded then unloaded
CM = pile material factor = 1 (drilled
augered) 09 (prestressed or precast
concrete) 08 (timber) and 07
(steel) and CK = pile installation factor
= 09 (bored or augered) 10 (low
displacement eg H-pile or open-end
pipe) and 11 (driven high
displacement eg prestressed
concrete closed-end pipe)
As the beta approach applies to all types of soils (gravels sands silts and clays) and more or
less has remained unchanged since its advent circa 1970 it has been selected for further
discussion herein In the direct form for calculation of unit pile side friction the expression is
given by
119874119862119877119904119894119899120601 prime prime 119891119901 = 119862119872 ∙ 119862119870 ∙ (1 minus 119904119894119899120601prime) middot ∙ 119905119886119899120601prime ∙ 120590119907119900 C5
C-10
where CM is a soil-pile interface friction coefficient and CK = pile installation factor Values of CM are in
the range 07 lt CM lt 10 depending upon pile material (steel wood concrete) and ranges for CK vary
09 lt CK lt 11 depending upon method of installation (auger drill driven) as detailed in Table 2
The value of K0 is limited to the passive stress coefficient which for the simple Rankine case is given by
KP = (1+sin ϕ) (1 - sin ϕ ) Thus there is a maximum value of overconsolidation ratio for which
Equation C5 applies given by OCRlimit = [(1+sin ϕ ) (1 - sin ϕ )2] (1sinϕ)
The corresponding graph in Figure C3 shows the relationship for β in terms of ϕ and OCR
Figure C3 Pile side friction coefficient β in terms of effective friction angle and overconsolidation ratio
C113 Pile Unit End Bearing
C1131 Theoretical Considerations
The evaluation of the end bearing resistance of pile foundations is commonly determined using
limit plasticity theory where
prime Drained loading = C6 119902119887 119873119902 middot 120590119907119900
C-11
Undrained loading 119902119887 = 119873119888 middot 119904119906 C7
where the value of σvo is calculated at the depth z = L and the value of su is the average undrained shear
strength from z = L to a depth z = L + d beneath the pile tip For a circular pile the limit plasticity solution
for undrained loading gives (Vesic 1977) a value Nc = 933 while a deep strip foundation would employ a
value of Nc = 824 For drained loading the expression for Nq for a deep foundation is given by (Vesic
1977)
)]arctan()sin1(tan21[)](tan1[sin1
sin1)tanexp( 2 BLABNq
C8
where A and B are the pile plan dimensions (for a circular pile A = B) and L = pile length For a
circular pile this can be approximated by
asymp 077(120601prime75deg) 119873119902 C9
over a range of effective friction angles 20deg le ϕ le 45deg
C1132 Practical Considerations
For drained loading of sands the full calculated end bearing capacity will never be realized because it
would require the pile to move a distance equal to its diameter This is beyond practical use and
therefore the limit plasticity solutions must be clipped to a fraction of the calculated value Another
reason for using a reduced end-bearing capacity is due to strain incompatibility since the side capacity is
mobilized early while the end-bearing is engaged much later So to have compatible values of Qs and
Qb the qb must be reduced using the following guidelines (Randolph 2003 Mayne 2007)
prime prime C10 119902119887 = 119873119902 ∙ 120590119907119900 ∙ 119891119909
where fx = strain incompatibility factor
fx = 010 for drilled shafts augered cast-in-place and bored piles
fx = 020 for driven low displacement piles (opened-ended steel pipe and H-piles)
fx = 030 for driven high-displacement piles (ie solid piles PSC and closed-ended pipe)
C-12
For undrained loading of circular piles in compression no reduction of end-bearing is necessary
therefore
119902119887 = 933 middot 119904119906 C11
C12 Direct CPT Methods for Pile Capacity
In the direct CPT method the penetrometer readings are scaled directly via specified algorithms to
obtain the pile unit side friction and end-bearing as depicted in Figure C4 As many as 40 different
direct CPT methods have been developed over the past five decades as summarized by Niazi amp Mayne
(2013) Starting circa 1970 many of these early methods relied on hand-recorded data from field
mechanical CPTs where only qc data were obtained at 20 cm intervals or later with mechanical readings
of both qc and fs using special sets of inner and outer rods that recorded vertical load for tip and loads
for tip plus sleeve in alternating increments Also early pile load test data were often obtained from
top-down measurements of load-displacement
C-13
Undrained qb = Nc su
Qside = S (fp DAs)
QTotal = Qs + Qb - Wp
fp = cmck Ko svorsquo tanrsquo
Qbase = qb Ab
Drained qb = Nq svorsquo
Method Two Rational or ldquoIndirectrdquo Method
Method OneldquoDirectrdquo CPT Method
(Scaled Pile)
fp = fctn (soil type pile
type qt fs and u2)
qb = fctn (soil type qt-u2)
qb = unit end bearing
unit sidefriction fp
OCR su Ko t DR rsquo
AXIAL PILE CAPACITY FROM CPT READINGS
Figure C4 Concept of Direct versus Rational Method for CPT evaluation of axial pile capacity
Beginning in the mid-1990s as the modern electric piezocone (CPTu) was implemented the newer
equipment offered improved data and better resolution because of the additional reading of porewater
pressures and correction of raw measured qc to total resistance qt In addition the use of electronic
digital data collection and field computers proved superior in field measurements and recordings As a
consequence several reliable CPT methods for axial pile capacity have been developed Moreover
parallel improvements in full-scale pile load testing occurred and now include modern strain gage
instrumentation digital data recording and automated testing procedures Also the testing can be
single direction (compression or tension) or bi-directional as in the Osterberg cell
C-14
Table C3 provides a selection of recent direct CPT methods that have become available over the
past two decades A number of these (namely ICP NGI UWA and Fugro) were funded by the
offshore industry because of the growth of oil amp gas reserves and windfarm installations thus
necessitating increased concerns on risk probability and reliability in the site investigations for
offshore platforms and design of large driven monopile foundations These direct CPT methods
are often statistically based on large datasets compiled from full-scale load tests made
worldwide (eg Schneider et al 2008)
Table C3 Selection of Direct CPT Methods for Axial Pile Capacity
Method Pile Types Soil Types References CPT
data
Additional
parameters
needed
Unicone Method all types Sands silts
clays
Eslami and
Fellenius
(1997)
Fellenius
(2009)
qt fs u2
KTRI = Kajima
Technical
Research
Institute
all types Sands
mixed clays
Takesue et al
(1998)
fs and u2 No guidance given
on end bearing
resistance
Imperial College
Procedure (ICP)
OE and CE Sands Chow et al
(1997 PhD)
Jardine et al
(2005)
qt Interface friction
angle (δ) from
ring shear tests
Correlated to
mean grain size
(D50)
C-15
OE and CE Clays Jardine et al
(2005)
qt Interface friction
angle (δ)
correlated to PI
Table C3 Continued
NGI Method
(Norwegian
Geotechnical
Institute)
OE and CE Sands Clausen et al
(2005)
qt DR from qt
Clays a Alpha
method
(Karlsrud et al
2005 2012)
b Beta
method
(Karlsrud
2012)
qt for su
qt for
OCR
PI = plasticity
index ()
PI = plasticity
index ()
Fugro Method OE and CE Sands Kolk et al
(2005)
qt
Clays Van Dijk and
Kolk (2011)
qt
OE and CE Sands Lehane et al
(2005)
qt IFR = infilling ratio
related to amount
of plugging
C-16
Purdue LFRD Drilled
shafts
Sands Basu amp
Salgado
(2012)
qt LRFD = load
resistance
factored design
Enhanced Various Sands silts Niazi and qt u2 Based on 330 load
Unicone Method clays and Mayne (2015 and fs tests including
mixed soils 2016) bored augered
jacked and driven
piles
UWA (Univ of
Western
Ra = pile
roughness
Australia) Clays Lehane et al
(2012)
qt Interface friction
angle (δ) from
ring shear tests
and also method
without δ
HKU Method
(Hong Kong
Univ)
OE and CE
(end
resistance
only)
Sands Yu and Yang
(2012)
qt End bearing only
PLR = plug length
ratio
Note PLR can be
estimated from
OE inner diam
Table C3 Continued
Note previously called Marine Technical Directorate (MTD)
C-17
Many of the offshore CPT methods relate to driven piles either in sand or clay as that is
common deep foundation for those purposes Of particular interest are the Unicone and
Modified Unicone Methods since they use all three readings of the piezocone (qt fs and u2)
and address a variety of pile foundation types
C13 Modified UniCone Method
The modified UniCone Method was developed based on a total 330 pile load tests which were
associated with SCPTu data during their site investigations (Niazi and Mayne 2015 2016) This
represents a threefold increase over the original UniCone database that was built upon data
from 106 pile load tests (Eslami amp Fellenius 1997)
For the original UniCone algorithms use is made of the effective cone resistance (qE)
119902119864 = 119902119905 minus 1199062 C12
and a chart of qE vs fs provided an approximate soil classification in five distinct groups as shown by
Figure C5 Later in the modified approach a better delineation of the larger dataset gave soil
subclassifications as indicated by Figure C6
C-18
Figure C5 Soil behavior type using CPT via original UniCone
Figure C6 Soil behavior type using CPT via Modified UniCone
C-19
In the modified approach the 9-zone normalized soil behavioral type (SBTn) is ascertained using CPT
data in conjunction with the Robertson (2009) charts as shown in Figure C7 As discussed in Appendix
A and B the SBTn method uses the normalized cone resistance (Qtn) normalized sleeve friction (Fr())
and CPT material index Ic This permits a much wider range in the calculated pile side friction because fp
is related as a continuous curve with Ic rather than only 5 values that are assigned in the original
scheme
The pile unit side friction (fp) is obtained from qE and the CPT material index Ic using the following
expression at each elevation along the sides of the pile
10(0732 middot 119868119888 minus 3605) fp middot middot C13 = 119902119864 middot 120579119875119879 middot 120579119879119862 120579119877119860119879119864
C-20
Figure C7 Nine-zone soil behavioral soil type using normalized piezocone parameters
(after Robertson 2009 Mayne 2014)
where θPT = coefficient for pile type (θPT = 084 for bored piles 102 for jacked piles 113 for driven
piles) θTC = coefficient for loading direction (θTC = 111 for compression and 085 for tension) and θRATE
= rate coefficient applied to soils in SBT zones 1 through 7 (θRATE = 109 for constant rate of penetration
tests and 097 for maintained load tests) Since CPT provides data at regular intervals of 2 cm to 5 cm
along the sides of the pile the average fp from z = 0 to z = L can be used directly in Equation 1 to obtain
the shaft capacity Qs
C-21
The pile end bearing resistance is obtained from
middot 10(0325middot 119868119888minus1218) 119902119887 = 119902119864 C14
where qE is averaged in the vicinity of the pile tip Figure C8 shows the Modified Unicone Method in
graphical format For sensitive clays of zone 1 please see additional discussions by Niazi amp Mayne
(2016)
Figure C8 Summary of Modified UniCone Method for direct CPT assessment of axial pile
capacity
C-22
C14 Axial Pile Displacement
The movement of pile foundations can be assessed using elastic continuum theory (Poulos amp
Davis 1980 Randolph 2003 Mayne amp Niazi 2017) These relationships have been developed
using finite element analyses boundary elements and analytical closed-form solutions For the
latter approach the top displacement of a rigid pile subjected to a vertical force is shown in
Figure C9 This gives the movement at the top of a pile subjected to either compression or
tension (uplift) loading Also the percentage of axial load transferred to the pile toe is
determined For piles extending through various soil layers the elastic solution can be
implemented by using a set of stacked pile segments each with its own stiffness as
represented by a soil Youngs modulus
For pile groups the use of computer software is recommended Several available programs can handle
pile groups under axial and lateral moment loading such as DEFPIG (Univ Sydney) and PIGLET (Univ
Western Australia) A full listing of pile foundation software is given at the Geotechnical amp
GeoEnvironmental Service Directory wwwggsdcom
Ps
= Sh
aft
Load
sL
t
tEd
IPw
Load Transfer to Base
)]1)((5ln[
)(
)1(1
1
2 vdL
dLFI
E
ECT
211
IF
P
P CT
t
b
Base Load = Pb
Randolph Elastic Solution
Ground Surface
Top Load Pt = Ps + Pb
Rigid Pile Response
wt = displacementd = diameter
L = length
Es = Elastic Soil Modulus
Es0 (at z = 0) surface
EsM (at z = L2)mid-length
EsL (at z = L)full length
Load Transfer to Shaft
E = EsMEsL = Gibson parameterE = 1 for homogeneous case (constant E)E = 05 for pure Gibson case (Es0 = 0)
where FCT = load direction (= 1 compression 0 tension)
21
IF
P
P CT
t
b
z = depth
= Poissons ratio
Tension
Compression
C-23
Figure C9 Elastic continuum solution for axial pile response under compression and tension
loading
C-24
C15 Acknowledgements
The author appreciates the help of Fernando Illingworth of PonsEcuador Meena Viswanth of
GeoSyntecAtlanta and Dr Marco Uzielli of GeoRiskItaly in helping to collect and process the data
Funding support was provided by David Woeller of ConeTec Richmond BC
C-25
- Structure Bookmarks
-
- CONE PENETRATION TEST DESIGN GUIDE FOR STATE GEOTECHNICAL ENGINEERS
- TABLE OF CONTENTS
- LIST OF TABLES
- CHAPTER 1 INTRODUCTION
- CHAPTER 2 DIRECT CPT METHOD FOR SOIL CHARACTERIZATION
- CHAPTER 3 DIRECT CPT METHOD FOR SHALLOW FOUNDATIONS
- CHAPTER 4 DIRECT CPT METHOD FOR DEEP FOUNDATIONS
- REFERENCES
- APPENDIX A
- APPENDIX B
- APPENDIX C
-